Technical Field
The present invention relates to a swirl stabilized gas turbine oxy-combustor, a method for oxy-combustion of a fuel using the swirl stabilized gas turbine oxy-combustor, and a monolith structure ion transport membrane reactor.
Description of the Related Art
Global climate change is one of the greatest challenges in the 21st century. The greenhouse gas making the largest contribution to global climate change from human activities is carbon dioxide (CO2). CO2 emissions from the fossil fuel-based large power plants are of main concern as they are the largest sources of CO2 in the coming decades. International Energy Agency forecasts have indicated that some 38 percentage of the world's electricity will still be generated from coal by 2020. See Priddle R., IEA World Energy Outlook, Paris, 1998, incorporated herein by reference in its entirety. For decreasing greenhouse gas (mainly CO2) emissions, several approaches have been evaluated and reviewed for capturing CO2 in the utility industry, namely Carbon Capture and Storage technology (CCS), including pre-combustion capture, oxyfuel combustion, and post-combustion capture. As a promising CCS technology, oxyfuel combustion can be used in existing and new power plants. See Buhre B. J. P., Elliott L. K., Sheng C. D., Gupta R. P., Wall T. F., Oxy-fuel combustion technology for coal-fired power generation, Prog. Energy Combust. Sci. 2005 Vol. 31, 283-307, incorporated herein by reference in its entirety. Compared to conventional air combustion, oxyfuel combustion shows different characteristics of heat transfer, ignition, char burnout as well as NOx emission. See Wall T. F., Combustion processes for carbon capture. Proceedings of the Combustion Institute, 2007, Vol. 3 1, 31-47, incorporated herein by reference in its entirety. During oxyfuel combustion, a combination of oxygen and recycled flue gases are used for combustion of the fuel. The exhaust gases consisting mainly of CO2 and H2O generated with a concentration of CO2 ready for sequestration. The recycled flue gases used to control flame temperature and make up the volume of the missing N2 to ensure there is enough gas to carry the heat through the boiler. CO2 capture and storage by the current technically viable options (post-combustion capture, pre-combustion capture and oxyfuel combustion) will impose a 7-10% efficiency penalty on the power generation process. The major contributors to this efficiency penalty are oxygen production and CO2 compression. Cryogenic air separation systems use intricately designed multi-stage distillation processes at low temperature and require a significant amount of electrical power, mainly for air compression. See Pfaff I., Kather A., Comparative thermodynamic analysis and integration issues of CCS steam power plants based on oxy-combustion with cryogenic or membrane based air separation, Energy Procedia. 2009, 1(1):495-502, incorporated herein by reference in its entirety. Typical electrical power requirements range from 160 kWh, to 270 kWh per ton of O2 with a commonly cited approximate value of 200 kWh per ton, depending on the desired purity. See Darde A., Prabhakar R., Tranier J. P., Perrin N., Air separation and flue gas compression and purification units for oxy-coal combustion systems, Energy Procedia. 2009, 1(1):527, 34 and Amann J. M., Kanniche M., Bouallou C., Natural gas combined cycle power plant modified into an O2/CO2 cycle for CO2 capture, Energy Conversion and Management 2009, 50(3):510-21, each incorporated herein by reference in their entirety. Despite their complexity, commercial cryogenic units achieve low Second Law efficiencies in the range of 15-24%, resulting in a First Law efficiency loss of up to 8.5% points compared to a typical NGCC (Natural gas combined cycle) without CCS. See Simpson A. P., Simon A. J., Second law comparison of oxy-fuel combustion and post-combustion carbon dioxide separation, Energy Conversion and Management. 2007, 48(11): 3034-45 and Kakaras E., Koumanakos A., Doukelis A., Giannakopoulos D., Vorrias I., Oxyfuel boiler design in a lignite-fired power plant, Fuel 2007, 86(14): 2144-50, each incorporated herein by reference in their entirety. Clearly, the thermodynamic and economic penalties incurred by the use of a cryogenic air separation process could easily offset any advantages gained by using Oxy-fuel CCS, prompting many researchers to investigate the use of alternative air separation systems.
Energy production from fossil fuel combustion results in the emission of greenhouse gases, the dominant contributor being CO2. Public awareness and legislation have led to a policy of reduction of greenhouse gas emissions in most economically well-developed countries, with the regulations partially driven by (international) initiatives such as the Kyoto protocol and the Inter-governmental Panel on Climate Change. See IPCC, Intergovernmental Panel on Climate Change, available at http://www.ipcc.ch.2004, Access date: Oct. 27, 2004, incorporated herein by reference in its entirety. Greenhouse gas emissions from energy production can be reduced by the use of alternative energy sources such as nuclear power and renewable energy. Renewable energy sources are increasingly used, however, until these sources can reliably produce significant amounts of energy, the immediate energy demand is likely to be met by conventional fossil fuel combustion, as indicated by energy policies and use projections. To reduce greenhouse gas emissions from fossil fuel-fired power generation, several possibilities exist including improving efficiency of power plants, introduction of combined cycles with generation by gas and steam turbines, which can achieve high thermal efficiencies, replacement of fossil fuels with renewable resources, substitution of coal by natural gas (having a lower carbon content), and CO2 capture and storage (called CCS). Incremental reduction of greenhouse gas emissions can be achieved by the first four options, however, to make a step-change reduction in emissions, the CO2 generated from combustion needs to be captured and stored (or sequestered).
All technologies include compression of the CO2 product to a supercritical state, typically 10 MPa, prior to transport and geological storage at a depth (and thereby pressure) retaining this state. FIG. 1 is an illustration of the three capture processes of CO2. Three main methods can be envisaged for the capture of CO2 with these three systems as shown in FIG. 1: (1) pre-combustion capture: to capture CO2 in a synthesis gas after conversion of CO into CO2; (2) post-combustion capture: to capture CO2 in the exhaust gases once the fuel has been fully burned with air; (3) capture in oxycombustion: consisting of combustion in oxygen with recycling of exhaust gases (therefore, composed mainly of CO2 and water) and purification of the CO2 flow, to eliminate incondensable gases.
FIG. 2 is an illustration of the efficiency of the different systems with and without CO2 capture. The literature shows (see FIG. 2) that the highest efficiency with capture is obtained for NGCC with post-combustion capture, with an efficiency of 50% compared with 60% without capture. The next highest is the oxy-combustion in PC, with an efficiency of almost 35% compared with 45% without capture, and then IGCC-Puertollano with pre-combustion capture with an efficiency of the order of 33.5% compared with 44% without capture. The lowest efficiency is obtained with post-combustion capture in PC, equal to 30% which is 15 points less than PC without capture (if MEA is used). See Kanniche M., Bonnivard R. G., Jaud P., Marcos J. V., Amann J. M., Bouallou C., Pre-combustion, post-combustion and oxy-combustion in thermal power plant for CO2 capture, Applied Thermal Engineering. 2010, 30, 53-62, incorporated herein by reference in its entirety. Therefore, recommended to only consider pre-combustion capture in IGCC, post-combustion in NGCC and oxy-combustion in PC.
The characteristics of oxy-fuel combustion with recycled flue gas differ with air combustion in several aspects primarily related to the higher CO2 levels and system effects due to the recirculated flow, including the following: (1) To attain a similar adiabatic flame temperature (AFT), the O2 proportion of the gases passing through the burner is higher, typically 30%, than that for air (of 21%), necessitating that about 60% of the flue gas is recycled. (2) The high proportions of CO2 and H2O in the furnace gases result in higher gas emissivities, so that similar radiative heat transfer for a boiler retrofitted to oxy-fuel will be attained when the O2 proportion of the gases passing through the burner is less than the 30% required for the same AFT. (3) The volume of gases flowing through the furnace is reduced somewhat, and the volume of flue gas (after recycling) is reduced by about 80%. (4) The density of the flue gas is increased, as the molecular weight of CO2 is 44, compared to 28 for N2. (5) Typically, when air-firing coal, 20% excess air is used. Oxy-fuel requires a percent excess O2 (defined as the O2 supplied in excess of that required for stoichiometric combustion of the coal supply) to achieve a similar O2 fraction in the flue gas as air firing, in the range of 3-5%. See Khare S., Wall T., Gupta R., Elliott L., Buhre B., The 30th International Technical Conference on Coal Utilisation and Fuel Systems 2005, Coal 9 Technology: Yesterday-Today-Tomorrow, incorporated herein by reference in its entirety. (6) Without removal in the recycle stream, species (including corrosive sulphur gases) have higher concentrations than in air firing. (7) As oxy-fuel combustion combined with sequestration must provide power to several significant unit operations, such as flue gas compression, that are not required in a conventional plant without sequestration, oxy-fuel combustion/sequestration is less efficient per unit of energy produced.
The combustion of fuel in a mixture of recirculated flue gas (RFG) and oxygen, however, presents new challenges to combustion specialists. Several experimental investigations with oxy-firing pulverized coal burners report that flame temperature and stability are strongly affected. See Croiset E., Thambimuthu K. V., NOx and SO2 emissions from O2/CO2 recycle coal combustion, Fuel. 2007, Vol. 80, 2117-2121 and Rohan S., Wall T., Sulphur impacts during pulverised coal combustion in oxyfuel technology for carbon capture and storage, Progress in Energy and Combustion Science 37 (2011) 69e88, each incorporated herein by reference in their entirety. This work focuses on the investigation of the oxy-combustion of methane to see the effect of CO2 recirculation on combustion characteristics. The substitution of N2 with CO2 in the oxidizer will lead to a reduction of the flame speed as reported by Zhu et al. See Zhu D. L., Egolfopoulos F. N., Law C. K., Propagation and extinction of stretched premixed flames, Symposium (International) on Combustion. 1998, Vol. 21, Issue 1, 1419-1426, incorporated herein by reference in its entirety. This causes poor combustion performance and a modified distribution of temperature and species in the combustion chamber.
Today, most of the researchers in the oxy-combustion field are working on combustion using mixed ionic and electronic conducting ceramic membranes. These membranes have received increasing interest from academia and industry. A major industrial effort is currently devoted to the development of the mixed-conducting ceramic membrane reactor technology for partial oxidation of hydrocarbons, in particular, partial oxidation of methane to syngas. See Dyer P. N., Richards R. E., Russek S. L., Taylor D. M., Ion transport membrane technology for oxygen separation and syngas production, Solid State Ionics. 2000, 134, 21 and Lin Y. S., Microporous and dense inorganic membranes: current status and prospective, Sep. Purif. Technol. 2001, 25, 39-55, each incorporated herein by reference in their entirety. The membrane materials studied most extensively are lanthanum cobaltite perovskite type ceramics. See Zeng Y., Lin Y. S., Swartz S. L., Perovskite type ceramic membranes: synthesis, oxygen permeation and membrane reactor performance for oxidative coupling of methane, J. Membrane Sci. 1998, 150, 87-98, incorporated herein by reference in its entirety. New mixed-conducting ceramic membrane materials developed recently for the membrane reactor applications include modified perovskite ceramics having general formula of SrFeCo0.5Ox, brownmillerite structured ceramic represented by Sr1.4 La0.6GaFeO3−δ, and ceramic-metal dual phase membranes, such as Sr0.2La0.8Fe0.69Co0.1 Cr0.2Mg0.01O3+50Ag/50Pd, and thin dual phase membranes consisting of the chemically stable yttria-stabilized zirconia (YSZ) and Pd phases. See Balachandran U., Kleefisch M. S., Kobylinski T. P., Morissette S. L., Pei S., Oxygen ion-conducting dense ceramic membranes, U.S. Patent 5. 1997, 639,437; Schwartz M., White J. H., Sammels A. F., Solid state oxygen anion and electron mediating membrane and catalytic membrane reactors containing them, U.S. Patent, 6. 2000, 033,632; Chen C. C., Prasad R., Gottzmann C. F., Solid electrolyte membrane with porous catalytically-enhancing constituents (Assigned to Praxair Technology), U.S. Patent 5. 1999, 938,822; and Kim J., Lin Y. S., Synthesis and oxygen permeation properties of thin YSZ/Pd composite membranes, AIChE J. 2000, 46, 1521, each incorporated herein by reference in their entirety. All these membranes are oxygen semi-permeable with high oxygen permeation flux when the membrane is exposed to air and a hydrocarbon such as methane.
In order to decrease CO2 emissions, many approaches have been evaluated in order to capture CO2, namely Carbon Capture and Storage technology (CCS). As a promising CCS technology, oxy-fuel combustion can be used in existing and new power plants. In oxy-combustion, a fuel is oxidized in a nearly nitrogen-free, diluted mixture such that the products consist mainly of CO2 and water vapor, enabling a relatively simple and inexpensive condensation separation process facilitating CO2 capture. There are two main approaches available to utilize the oxy-combustion technology, one of them is through the use of air separation units to separate O2 which will be used in the combustion process and the other application is the ion transport membrane (ITM) reactor technology.
The combustion of fuel in a mixture of recirculated flue gas (RFG) and oxygen, however, presents new challenges to combustion specialists. Several experimental investigations with oxy-firing pulverized coal burners reported that flame temperature and stability are strongly affected. See Croiset E., Thambimuthu K. V., NOx and SO2 emissions from O2/CO2 recycle coal combustion, Fuel. 2007, Vol. 80, 2117-2121, incorporated herein by reference in its entirety. The substitution of N2 with CO2 in the oxidizer will lead to a reduction of the flame speed as reported by Zhu et al. See Zhu D. L., Egolfopoulos F. N., Law C. K., Propagation and extinction of stretched premixed flames, Symposium (International) on Combustion. 1998, Vol. 21, Issue 1, 1419-1426, incorporated herein by reference in its entirety. This causes poor combustion performance and a modified distribution of temperature and species in the combustion chamber. Liu et al. have performed numerical investigations on the chemical effects of CO2. See Liu F., Guo H., Smallwood G., The chemical effect of CO2 replacement of N2 in air on the burning velocity of CH4 and H2 premixed flames, J. Combust. Flame. 2003, Vol. 133 (4), 495-497, incorporated herein by reference in its entirety. A comparison between numerical and experimental data showed that the decrease in burning velocity for the oxyfuel combustion cannot entirely be described by only considering the material properties of CO2. Anderson et al. have performed experiments on a 100 kW test unit which facilitates O2/CO2 combustion with real flue gases recycle. See Andersson K., Johnsson F., Flame and radiation characteristics of gas-fired O2/CO2 combustion, Fuel. 2007, Vol. 86, 656-668, incorporated herein by reference in its entirety. The tests comprise a reference test with air and two O2/CO2 test cases with different recycled feed gas mixture concentrations of O2 (OF 21 @ 21 vol. % O2, 79 vol. % CO2 and OF 27 @ 27 vol. % O2, 73 vol. % CO2). The results showed that the fuel burnout is delayed for the OF 21 case compared to air-fired conditions as a consequence of reduced temperature levels. Instead, the OF 27 case results in more similar combustion behavior as compared to the reference conditions in terms of in-flame temperature and gas concentration levels, but with significantly increased flame radiation intensity.
On the other hand, Teraoka et al. first demonstrated that La1−xSrxCo1−yFeyO3−δ oxide-based perovskite-type ceramic membranes have appreciably high oxygen permeation fluxes at high temperatures. See Teraoka, Y., Zhang H. M., Furukawa S., Yamazoe N., Oxygen permeation though perovskite-type oxides, Chem. Lett. 1985, 1743, incorporated herein by reference in its entirety. Following Teraoka et al.'s work, many researchers studied the La—Sr—Co—Fe series as oxygen permeable membranes. See Kruidhof H., Bouwmeester H. J. M., Doom R. H. E., Burggraaf A. J., Influence of order-disorder transitions on oxygen permeability through selected nonstoichiometric perovskite-type oxides, Solid State Ionics. 1993, 3B-65B, 816; Qiu L., Lee T. H., Liu L. M., Yang Y. L., Jacobson A. J., Oxygen permeation studies of SrCo0.8Feo0.2O3−δ, Solid Sate Ionics. 1995, 76, 321; Stevenson J. W., Armstrong T. R., Carmeim R. D., Pederson L. R, Weber L. R., Electrochemical properties of mixed conducting perovskite La1−xMCo1−yFeyO3−δ (M=Sr, Ba, Ca), J. Electrochem. Soc. 1996, 143, 2722; Tsai C. Y., Dixon A. G., Ma Y. H., Moser W. R., Pascucci M. R., Dense perovskite La1−xMxCo1−yFeyO3−δ (M=Sr, Ba, Ca) membrane synthesis, application, and characterization, J. Am. Ceram. 1998, Soc. 81, 1437; Xu S. J., Thomson W., Stability of La0.6Sr0.4Co0.2Fe0.8O3−δ perovskite membranes in reducing and nonreducing environments, J. Ind. Eng. Chem. Res. 1998, 37, 1290; Elshof T. J. E., Bouwmeester H. J. M., Verweij H., Oxidative coupling of methane in a mixed conducting perovskite membrane reactor, Appl. Catal. 1995, A 130, 195; Gu X. H., Jin W. Q., Chen C. L, Xu N. P., Shi J., Ma Y. H., YSZ—SrCo0.4Fe0.6O3−δ elta membranes for the partial oxidation of methane to syngas, AIChE J. 2002, 48, 2051-2060; Zhang K., Yang Y. L., Ponnusamy D., Jacobson A. J., Salama K., Effect of microstructure on oxygen permeation in SrCo0.8Fe0.2O3−δ, J. Mater. 1999, Sci. 34, 1367, each incorporated herein by reference in their entirety. The oxygen ionic transference number obtained was in the range of 10−5 to 10−3 depending on the temperature and ambient atmosphere, and the activation energy for ionic transport was 64-125 kJ/mol. Inconsistencies (up to one order of magnitude) were frequently reported on the oxygen permeation rates even with nominally identical membrane materials. For example, Teraoka et al. reported the oxygen permeation flux as high as 2.31×10−6 mol/cm2s at 850° C. for the SrCo0.8Fe0.2O3δ membrane; while the oxygen permeation fluxes for SrCo0.8Feo0.2O3−δ membrane of the same thickness from Kruidhof et al. and Qiu et al. are 1.8×10−7 and 6.3×10−7 mol/cm2s, respectively, under the same experimental conditions. In membrane reactor applications in which the membrane is, respectively, exposed to air and methane, the difference in the oxygen permeation fluxes reported by different research groups could be as large as two orders of magnitudes even for membranes of same or similar compositions.
Experimental and numerical investigations of an atmospheric diffusion oxy-combustion flame in a gas turbine model combustor are disclosed herein. Oxycombustion and emission characterization, flame stabilization and oxy-combustion model validation analyses are the main goals of the present disclosure. The combustor is fuelled with CH4 and a mixture of CO2 and O2 as oxidizer. Wide ranges of different operating parameters are considered including equivalence ratio, percentage of O2/CO2 in the oxidizer mixture, and fuel volume flow rate. Stability of the oxy-combustion diffusion flame is checked experimentally and numerically. The minimum permissible percent of O2 in the oxidizer mixture required in order to get a stable flame is determined. Visualizations of the flame at the above mentioned operating conditions have been carried out experimentally and compared with the numerical results. The flames have been characterized in details by measuring the exhaust gas temperatures and comparing them with those from the numerical model. Flame zone also has been characterized in details by plotting the axial and radial temperatures, species concentrations and flow velocities. A modified two-step oxy-combustion reaction kinetics model for methane-oxygen combustion has been used in order to predict a clearer oxy-combustion characteristics and then more accurate numerical results in order to correctly validate the numerical model using the experimental results.
Computational fluid dynamics (CFD) is becoming an important industrial tool for trouble-shooting. However, CFD modeling of industrial combustion applications is a computationally demanding task. See Andersen J., Rasmussen C. L., Giselsson T., Glarborg P., Global Combustion Mechanisms for Use in CFD Modeling under Oxy-Fuel Conditions, Energy Fuels, 2009, 23, 1379-1389, incorporated herein by reference in its entirety. It is necessary to apply simplified reaction mechanisms to reduce the computing cost and time; however the simplified schemes do not work as well under oxy-fuel combustion conditions like the case of combustion using air. For conducting the 2D analysis of the simple symmetric design ITM reactor, the modified two step finite rate oxy-combustion reaction kinetics model done by Andersen et. al. is used to predict oxy-combustion characteristics inside a stagnation flow ITM reactor.
Liu et al. have performed numerical investigations on the chemical effects of CO2. A comparison between numerical and experimental data showed that the decrease in burning velocity for the oxyfuel combustion cannot entirely be described by only considering the material properties of CO2. CO2 affects the combustion reactions especially by the reaction CO+OH→CO2+H. The competition of CO2 for H radical through the above reverse reaction with the single most important chain branching reaction H+O2→O+OH significantly reduces the concentrations of important radicals, i.e. O, H, and OH, leading to significant reduction of fuel burning rate. This hypothesis is supported by a comparison of the burning velocity of methane flames and hydrogen flames in a CO22/O2 gas mixture.
The influence of CO2 on the burning velocity of hydrogen flames is less significant because the concentration of hydrogen radicals is much higher. The chemical effect of CO2 significantly reduces the burning velocity of methane, where by the relative importance of this chemical effect increases with increasing CO2 concentration in the oxidizing mixture.
During oxy-fuel combustion, the amount of NOx exhausted from the system can be reduced to less than one-third of that with combustion in air. See Jyh-Cherng C., Zhen-Shu L., Jian-Sheng H., Emission characteristics of coal combustion in different O2/N2, O2/CO2 and O2/RFG atmosphere, Journal of Hazardous Materials 142 (2007) 266-271 and Kimura K., Omata K., Kiga T., Takano S., Shikisima S., Characteristics of pulverized coal combustion in O2/CO2 mixtures for CO2 recovery, Energy Convers Manage 1995; 36: 805-808, each incorporated herein by reference in their entirety. The NOx reduction is thought to be the result of several mechanisms: Decrease of thermal NOx due to the very low concentration of N2 from air in the combustor, the reduction of recycled NOx as it is reburnt in the volatile matter release region of the flame, and the reaction between recycled NOx and char. Okazaki and Ando used a bench-scale reactor to examine the effects of the latter two factors during oxy-fuel combustion with an O2 concentration of 21% (i.e., recycling ratio as high as 80%) at a maximum flame dominant mechanism for the reduction in NOx emissions. See Okazaki K., Ando T., NOx reduction mechanism in coal combustion with recycled CO2, Energy, 22 (1997) 207-215, incorporated herein by reference in its entirety. They estimated that more than 50% of the recycled NOx was reduced when 80% of the flue is recycled. It has also been found that oxy-fuel combustion can decrease the SO2 emissions compared to that in air combustion. See Hu Y., Naito S., Kobayashi N., Hasatani M., CO2, NOx and SO2 emissions from the combustion of coal with high oxygen concentration gases, Fuel. 2000, 79, 1925-1932, incorporated herein by reference in its entirety.
The concept of oxy-combustion involves the burning of fuel in pure oxygen in addition to some recycled flow gases or steam in order to control the flame temperature. The aim is to obtain a carbon dioxide-rich stream that is ready for sequestration, after removing water vapor and other impurities. Various oxy-combustion systems have been introduced in the literature. See Buhre B. J. P., Elliott L. K., Sheng C. D., Gupta R. P., Wall T. F., Oxy-fuel combustion technology for coal-fired power generation, Prog. Energy Combust. Sci. 2005 Vol. 31, 283-307; Seepana S., Jayanti S., In: ASME international mechanical engineering congress and exposition, Boston, Mass., 2008, 435-444; Seepana S., Jayanti S., Steam-moderated oxy-fuel combustion, Energy Conversion and Management, 2010, 51, 1981-1988; Hong J., Chaudhry G., Brisson J., Field R., Gazzino M., Ghoniem A., Analysis of Oxy-Fuel Combustion Power Cycle Utilizing a Pressurized Coal Combustor, Energy. 2009, 34:1332-1340; and Andersen R., MacAdam S., Viteri F., Davies D., Downs J., Paliszewski A., In: Proceedings of ASME Turbo Expo, 2008, Berlin, Germany, each incorporated herein by reference in their entirety. The first version is the atmospheric pressure oxy-combustion system where part of the flue gases is recycled in order to control the flame temperatures. There is another alternative to using recycled flue gases is to inject steam in order to control the flame temperature. To further increase the performance of these systems, pressurized systems have been proposed for both systems oxy-combustion with recycled flue gases and oxy-syngas combustion in combination with solid fuel gasification technology. See Zheng L., Pomalis R., Clements B., Herage T., In: The 35th international technical conference on clean coal & fuel systems, 2010, Clearwater, Fla.; Fassbender A., Henry R., Tao L., AEA Report, AEA Grant Number. 2009, -AEA 07-014; Hong J., Thesis: Cambridge. 2009, Massachusetts Institute of Technology; and Hong J., Field R., Gazzino M., Ghoniem A., Operating Pressure Dependence of the Pressurized Oxy-Fuel Combustion Power Cycle, Energy 2010, 35, 5391-5399, each incorporated herein by reference in their entirety. Ion transport membrane reactor technology can also be applied and it is discussed in details later on. FIG. 3 is an illustration of the atmospheric oxy-coal combustion system with flue gas recycle proposed for carbon capture in coal power plants. The atmospheric oxy-coal combustion system shown in FIG. 3 was introduced as a short-term solution to retrofit existing coal-fired power plant to include the option of carbon capture and sequestration. The additional required equipments as compared with air-fired systems are considered herein.
When retrofitting existing power plants to be used with oxy-combustion, the system uses the same equipments used in the conventional combustion in addition to an ASU used to produce an oxygen rich stream for combustion. Currently, the only ASU technology that can meet the volume and purity demand of a large scale coal-fired utility boiler is based on cryogenic distillation. Air is compressed, cooled and cleaned prior to being introduced into the distillation column to separate air into an oxygen-rich stream and a nitrogen-rich stream. See Chen L., Zheng S., Yong, Ghoniem A., Modeling the slag behavior in three dimensional CFD simulation of a vertically-oriented oxy-coal combustion, Progress in Energy and Combustion Science. 2012, 38156-214, incorporated herein by reference in its entirety. Cryogenic air separation is consuming about 0.24 kWh/kg O2 with 95% oxygen purity. See Haslbeck J., Capicotto P., Juehn N., Lewis E., Rutkowski M., Woods M., et al., In: Bituminous coal to electricity, Vol. 1. Washington D.C. 2007, DOE/NETL-1291 and Haslbeck J., Capicotto P., Juehn N., Lewis E., Rutkowski M., Woods M., et al., In: Bituminous coal to electricity, Vol. 1. Washington D.C. 2007, DOE/NETL-1291, each incorporated herein by reference in their entirety. The oxygen purity requirement for oxy-coal combustion (85-98%) is lower than that needed in the process industry (99.5-99.6%). See Darde A., Prabhakar R, Tranier J. P., Perrin N., Air separation and flue gas compression and purification units for oxy-coal combustion systems, Energy Procedia. 2009, 1(1):527, 5534, incorporated herein by reference in its entirety. The ASU can consume more than 15% of the gross power output. See Andersson K., Johnsson F., Process evaluation of an 865 MWe lignite fired O2/CO2 power plant, Energy Conversion and Management. 2006, 47:3487-98; Okawa M., Kimura N., Kiga T., Takano S., Arai K., Kato M., Trial design for a CO2 recovery power plant by burning pulverized coal in O2/CO2, Energy Conversion and Management. 1997, 38:S123-7; and Varagani R., Chatel F., Pranda P., Rostam M., Lu Y., Bose A., In: The 4th annual conference on carbon sequestration, 2005, Alexandria, Va., each incorporated herein by reference in their entirety.
A carbon purification unit consists of gas cleanup units in order to remove water and other gases from the flue gas before being compressed for the sequestration process. Because oxy-combustion is compatible with retrofits, selective catalytic reduction (SCR), electrostatic precipitator (ESP) and flue gas desulphurization (FGD) are typically retained as means of NOx, particulate matter and SOx removal from the flue gases. These pollutants control devices are also suitable for use in conjunction with amine-type absorbents for post-combustion capture plants.
After the removal of acid gases such as SOx and NOx, non-condensable N2, O2, and Ar should also be purged using a non-condensable gas purification unit. This unit is made of multi-stage compression units with inter-stage cooling in order to separate out the inert gases.
Recycled flue gases are required for replacement of nitrogen in order to control the combustion temperature. These flue gases can be recycled at different locations downstream of the economizer in the form of wet or dry recycles. since SO2 concentration in the flue gas may accumulate due to flue gas recycle, resulting in two or three times higher concentration than in conventional air-firing systems, the primary recycle has to be at least partially desulphurized for medium and high sulfur coal, to avid corrosion in the coal mill and flue gas pipes.
Capture of CO2 from large point sources such as power plants with subsequent geological storage offers the possibility of a significant and relatively quick response to climate change at a reasonable cost. Successful commercialization of such technology could therefore provide a transition to a future during which energy production from non-fossil energy sources can grow over time. At present, there are no power plants with CO2capture available on a commercial scale, but long time aquifer storage is currently applied and evaluated in the North Sea and show promising results. See Torp T. A., Gale J., Demostrating storage of CO2 in geological reservoirs: the Sleipner and SACS projects, Energy. 2004, 29, 1361, 1369, incorporated herein by reference in its entirety. The CO2 could also be stored in connection to enhanced oil recovery (EOR). Such storage has been closely monitored in the Weyburn project in Canada, see e.g. Emberly S., Hutcheon I., Shevalier M., Durocher K., Gunter W. D., Perkins E. H., Geochemical monitoring of fluid-rock interaction and CO2 storage at the Weyburn CO2-injection enhanced oil recovery site, Energy. 2004, 29:1393, 1401, incorporated herein by reference in its entirety. The highest cost is however on the capture side and to reduce the specific costs for capture different concepts are discussed. To recover and store carbon dioxide from flue gases of fossil fuel power plants, processes based on oxy-combustion appear to be promising. Concept of the technology is the combustion with commercially pure oxygen to achieve high CO2 concentrations in the flue gases for the final CO2 separation. The required oxygen is supplied by an air separation unit where the nitrogen is separated from the air. A great portion of the flue gases has to be recycled to substitute the removed nitrogen. This measure is inevitable to maintain the temperature level in the combustion chamber and in particular not to increase the heat transferred to the membrane walls of the steam generator which is limited by material parameters. See Pfaff I., Kather A., Comparative Thermodynamic Analysis and Integration Issues of CCS Steam Power Plants Based on Oxy-Combustion with Cryogenic or Membrane Based Air Separation, Energy Procedia. 2009, 1, 495-502, incorporated herein by reference in its entirety.
In the past decades, intense research efforts have been directed to the development and improvement of ceramic-based membranes for oxygen separation from air at high-temperature operations. Ceramic based membranes for oxygen separation systems can be categorized into pure oxygen conducting membranes and mixed ionic-electronic conducting membranes. Mixed ionic and electronic conducting ceramic membranes have received increasing interest from academia and industry. A major industrial effort is currently devoted to the development of the mixed-conducting ceramic membrane reactor technology for partial oxidation of hydrocarbons, in particular, partial oxidation of methane to syngas.
The required oxygen in this case is supplied by an air separation unit where the nitrogen is separated from the air. A great portion of the flue gases has to be recycled to substitute the removed nitrogen. A key component of the oxy-fuel process with high temperature membrane air separation unit (HTM-ASU), which is in the stage of development, is a dense membrane made of ceramic materials. These materials begin to conduct oxygen ions above a material dependent temperature (usually above 700° C.). Driving force for the mass transport is the differential oxygen partial pressure across the membrane, while the oxygen flux is enhanced with decreasing membrane thickness and rising temperature. As only oxygen permeates the membrane, 100% pure oxygen could be produced provided that air leakage within the membrane module is avoided. For further details regarding membrane materials, references are recommended. See Li K., Ceramic Membranes for Separation and Reaction. 2007, John Wiley & Sons, Ltd., West Sussex and Sirman J., Chapter 6 In: Nonporous Inorganic Membranes. 2006, WILEY-VCH Verlag GmbH & Co. K G, Weinheim, pp. 165-184, each incorporated herein by reference in their entirety. FIG. 4 is a schematic of an air separation unit based on high temperature membranes (exemplary flue gas swept). The basic idea of the HTM-ASU, as illustrated schematically in FIG. 4, is the elevation of the oxygen partial pressure on the air side with an air compressor. The partial pressure difference across the membrane can be further enhanced by lowering the oxygen partial pressure on the oxygen receiving side of the membrane by sweeping with flue gas, which contains only a small amount of oxygen. As temperatures at the compressor outlet are not sufficient to activate the membrane material's conduction mechanism, the air needs to be preheated with counter current oxygen enriched flue gas. To recover parts of the spent energy for compression, the oxygen depleted air is expanded in a turbine. As the off-gas leaves the HTM-ASU at still elevated temperatures, the heat can be recovered in the power plant cycle. The energy demand of the HTM-ASU is determined by the required high temperature heat. In addition, mechanical driving power is needed or produced depending on the ASU process parameter design.
As an option to get O2 required for combustion, Oxygen may be obtained via air separation units, e.g. cryogenic or membrane based processes. The combustion process takes place in a nitrogen free or low-nitrogen environment resulting in a flue gas composed mainly of CO2 and H2O, as well as a low concentration of impurities such as argon and oxygen. Therefore, a simplified flue gas processing by means of condensation of H2O to capture CO2, without using costly separation methods such as chemical absorption, can be possible. There are several proposed combined cycle concepts in oxy-fuel gas turbine processes with natural gas combustion in oxygen and CO2, for example, the O2/CO2 cycle, the COOLENERG cycle, the COOPERATE cycle, and the MATIANT cycle. See Kvamsdal H. M., Jordal K., Bolland O., A quantitative comparison of gas turbine cycles with CO2 capture, Energy 2007; 32:10-24; Bolland O., Mathieu P., Comparison of two CO2 removal options in combined cycle power plants, Energy Convers Manage 1998; 39(16-18):1653-63; Dillon D. J., Panesar R. S., Wall R. A., Allam R. J., White V., Gibbins J., et al., Oxycombustion processes for CO2 capture from advanced supercritical PF and NGCC power plant, In: Proceedings of the seventh international conference on greenhouse gas control technologies—GHGT7. Vancouver, Canada; September, 2004; Staicovici M. D., Further research zero CO2 emission power production: the COOLENERG process, Energy 2002; 27:831-844; Yantovski E. I., Stack downward zero emission fuel-fired power plants concept, Energy Convers Manage 1996; 37:867-877; Mathieu P., Nihart R., Sensitivity analysis of the MATIANT cycle, Energy Convers Manage 1999; 40:1687-700, each incorporated herein by reference in their entirety. These cycles belong to the group also known as Semi-Closed Oxy-Fuel Combustion Combined Cycles (SCOF-CC). Recent studies within the European Union funded research project ENCAP (Enhanced Capture of CO2) have concluded that SCOF-CC has good potential with limited technoeconomical hinders for realization. See The European Technology Platform for Zero Emission Fossil Fuel Power Plants (ZEP), The final report from working group 1 power plant and carbon dioxide capture; 13 Oct. 2006 and Sanz W., Jericha H., Bauer B., Göttlich E., Qualitative and quantitative comparison of two promising oxy-fuel power cycles for CO2 capture, Paper GT2007-27375, ASME Turbo Expo, Montreal, Canada; 2007, each incorporated herein by reference in their entirety. Besides SCOF-CC a number of other oxyfuel cycles using steam/water as working fluids have been proposed including the Graz cycle, and the Water cycle developed by Clean Energy Systems (CES). See Jericha H., Sanz W., Göttlich E., Design concept for large output graz cycle gas turbines, ASME Paper GT2006-90032, ASME Turbo Expo 2006, Barcelona, Spain; 2006 and Anderson R. E., MacAdam S., Viteri F., Davies D. O., Downs J. P., Paliszewski A., Adapting gas turbines to zero emission oxy-fuel power plants, Paper GT 2008-5 1377, ASME Turbo Expo, Berlin, Germany; 2008, each incorporated herein by reference in their entirety. These cycles may require high temperature turbines and new design for the turbomachinery. For oxy-fuel gas turbine cycles, researches have been focused on thermodynamic studies of system performance. The combustion behavior, e.g. the flame dynamics and reaction zone structures in the gas turbine combustors, is less addressed. From thermodynamic studies it has been shown that a small amount of trace species in the combustion products can have a great impact on the CO2 capture, storage and transportation. Li et al. demonstrated that the purification process of the flue gas stream of oxy-fuel combustion is highly influenced by the existence of impurities such as the small amount of N2 resulted from the air separation units and the remaining O2 in the flue gas due to incomplete combustion. See Li H., Yan J., Yan J., Anheden M., Impurity impacts on the purification process on oxy-fuel combustion based CO2 capture and storage system, Appl Energy 2009; 86:202-13, incorporated herein by reference in its entirety. The presence of non-condensable gases results in increased condensation duty for the recovery of the CO2. This in turn leads to lower system efficiency and increased cost for separation. To minimize the oxygen concentration in the flue gas and meanwhile achieve complete combustion of fuel, stoichiometric mixture is preferred in oxy-fuel combustion. CO2 and/or steam are used to control the flue gas temperature. Jericha and Gottlich outlined a burner and combustor configuration, in which fuel, oxygen and steam were supplied separately in different inlets. See Jericha H., Göttlich E., Conceptual design for an industrial prototype graz cycle power plant, Paper GT 2002-30118, ASME Turbo Expo, Amsterdam, Netherlands; 2002, incorporated herein by reference in its entirety. The steam was supplied through an annular outer swirler inlet to form a swirling flow motion to wrap the flames and to cool down the flue gases.
Such combustor configuration would likely generate rather high flame temperature locally in the reaction zones that will enhance the dissociation of H2O and CO2 and thus affect the composition of the flue gas such that the un-consumed oxygen can be high in the flue gas. To reduce the flame temperature and thereby the remaining oxygen in the flue gas it can be beneficial to premix the oxygen and CO2 or steam before injecting them to the combustor. There are several possibilities that need to be explored for example, different levels of premixing of the fuel/oxygen/steam/CO2 prior to their injection into the combustor, and different mixing patterns inside the combustor. The thermodynamic studies will give the same answer for the flue gas in the post flame zone if the inlet temperature, combustor pressure and the overall mass flows of fuel, oxygen, steam, and CO2 streams are kept the same. However, the flame dynamics and reaction zone structures are dependent on combustor configurations as they are dictated by the detailed inflow conditions for the fuel/oxygen/steam and CO2 supplies.
Optimal Supply of Oxygen and Diluent to Oxy Fuel Combustion:
To generate stable combustion in gas turbine combustion chambers with oxy-fuel combustion, certain minimal oxygen level in the oxidizer or elevated oxidizer temperature has to be maintained. The fundamental reason for this is the need to have sufficiently high temperature in the reaction zones for the chain reactions to proceed. Flame instability and poor burnout have been experienced when oxygen/CO2 are premixed and supplied together to the flame as the oxidizer. See Woycenko D., van de Kamp W., Roberts P., European Comission Joule II clean coal technology program 1992-1995, vol. II. Powder coal combustion projects final reports; 1997. ISBN:92-9-828-006-7, incorporated herein by reference in its entirety. For example, in the recent experiments of Heil et al. it was shown that poor burnout and lifted dark flames appeared when the oxygen mole fraction in the O2/CO2 stream was set to 21%; when the oxygen volume fraction was increased to 27% and 34%, full burnout and stable flames were obtained. See Heil P., Torporov D., Stadler H., Tschunko S., Forster M., Kneer R., Development of an oxycoal swirl burner operating at low O2 concentrations, Fuel. 2009, 88, 1269-1274, incorporated herein by reference in its entirety. In order to burn the fuel with lower oxygen level in the oxidizer (O2/CO2) stream the burner had to be modified to allow for recirculation of hot gases to the flame. To improve the recirculation, the inlet design of the oxidizer mixture should be modified in order to mix the hot burned gases in the flame zone with the incoming fresh cold gases in order to stabilize the flame.
In the study done by Kutne et.al., the burner considered was a modified version of a practical gas turbine combustor with an air blast nozzle for liquid fuels. See Kutne P., Kapadia B. K., Meier W., Aigner M, Experimental analysis of the combustion behavior of oxyfuel flames in a gas turbine model combustor, proceedings of the combustion institute.2010.doi: 10.1016/j.proci.2010.07.008 and Cao M., Eickhoff H., Joos F., Simon B., in: ASME Propulsion and Energetics, 70th Symposium, AGARD Conf. Proc. 422, 1987, 8.1, each incorporated herein by reference in their entirety. Co-swirling oxidizer mixture was supplied to the flame through a central nozzle and an annular nozzle. The radial swirlers consisted of 8 channels for the central nozzle and 12 channels for the annular nozzle. The overall flow field of the flames is characterized by a conically shaped inflow of fresh gas, an inner recirculation zone (IRZ) and outer recirculation zone (ORZ) as sketched in FIG. 5. FIG. 5 is a schematic of diagram of a combustion chamber. In the shear layer formed between the inflow and the IRZ, the mixing of hot combustion products with fresh gas leads to a continuous ignition and stabilization of the flame. Same idea of conically shaped inlet for the oxidizer mixture is applied to the running system of gas turbine model combustor in order to improve the mixing process and so stabilizing the flame. However, the aspect ratio between the diameter of the inlet nozzle and the combustor diameter plays an important role in calculating the amount of the oxidizer mixture that will be available for combustion and the amount that will escape between the flame zone and the reactor walls.
With high level oxygen in the oxidizer, the combustion products become hot and this may lead to high level of oxygen in the flue gas due to the dissociation reactions at high temperatures. There is an optimal window” of oxygen/diluent ratio in the oxidizer stream. See Liu C., Chen G., Sipöcz N., Assadi M., Bai X., Characteristics of oxy-fuel combustion in gas turbines, Applied Energy. 2012, 89, 387-394, incorporated herein by reference in its entirety. In the work done by Liu et al., they reported that the primary oxidizer which is supplied in the upstream through the dome of the combustion chamber should have minimal oxygen level of 24% under the oxidizer temperature 520 K condition. The reaction zones to have a suitable temperature when the flue gas enters to the turbines. This will cool down the combustion products generated in the primary reaction zones. Stable combustion and low turbine inlet temperature can be obtained simultaneously by adjusting the oxygen and CO2 supplies to the combustion chamber.
The stability of swirl stabilized oxyfuel/CH4 flames was studied in the work done by Peter Kutne et al. for O2 mole fractions of 20-40%, equivalence ratios of Φ=0.5-1 and thermal powers of 10-30 kW. However, attempts of operating the burner with <22% O2 were unsuccessful even with conditions of Φ=1 at 20 and 30 kW resulting in unstable operation and blow out.
Reactions Characteristics:
The oxyfuel combustion of coal in a steam turbine process is regarded as a possible way to use the oxyfuel process for CO2 reduction. Research on this field is very active with the outcome that the first demonstration plants are in operation, and the power generation industry is willing to invest in this technology. Another way of particular interest is the use of oxyfuel combustion in gas turbines. This process offers the possibility to use the same post combustion techniques as for the oxyfuel coal process, in combination with an efficient combined cycle process. Swirl flames are used extensively in practical combustion systems because they enable high energy conversion in a small volume and exhibit good ignition and stabilization behavior over a wide operating range. See Gupta A., Lilley D., Syred N., Swirl Flows, 1984, Abacus Press, Kent; Syred N., Chigier N. A., Beér J. M., Flame Stabilization in Recirculation Zones of Jets with Swirl, Thirteenth Symposium on Combustion, University of Utah, Salt Lake City, 1970. Thirteenth International Symposium on Combustion, Combust. Inst. 1971, 13, 617-624; Syred N., Beér J. M., Combustion in swirling flows: A review, Combust. Flame. 1974, 23, 143-201; and Weber R., Dugué J., Combustion accelerated swirling flows in high confinements, Prog. Energy Combust. Sci. 1992, 18, 349-367, each incorporated herein by reference in their entirety. In stationary gas turbine (GT) combustors, they are used mostly as premixed or partially premixed flames, and in aero engines, as diffusion flames. To reduce pollutant emissions, especially NOx, the flames are operated generally very lean. See Correa S. M., Power generation and aeropropulsion gas turbines: from combustion science to combustion technology, Combust. Inst. 1998, 27, 1793-1807; Lefebvre A. H., Gas Turbine Combustion, Taylor & Francis, Philadelphia, 1999; and Bauer H. J., New low emission strategies and combustor designs for civil aeroengine Applications, Prog. Comput. Fluid Dyn. 2004, 4, 130-142, each incorporated herein by reference in their entirety. Under these conditions, the flames tend to exhibit undesired instabilities, e.g., in the form of unsteady flame stabilization or thermoacoustic oscillations. The underlying mechanisms of the instabilities are based on the complex interaction between flow field, pressure, mixing, and chemical reactions, and are not well enough understood to date. Detailed measurements in full-scale combustors are hardly possible, and very expensive and numerical tools have not yet reached a sufficient level of confidence to solve the problems. A promising strategy lies therefore in the establishment of a laboratory-scale “standard combustor” with practical relevance and detailed, comprehensive measurements using nonintrusive techniques with high accuracy. The gained data set will be used for validation and measurement of numerical combustion simulation codes which then can be applied to simulate the behavior of technical combustors. Intrusive probe measurements are less suited for these applications as they disturb the local flow field and change the conditions for stabilization and for reaction—locally or even in general. See Meier W., Duan X. R., Weigand P., Lehmann B., Temperatur-Messungen in turbulenten Drallflammen: Thermoelemente im Vergleich zu Laser-Raman-Streuung, Gaswairme Int. 2004, 53, 153-158 and Stricker W., Kohse-Höinghaus K., Jeffries J., Applied Combustion Diagnostics, Taylor & Francis, New York, 2002, pp. 155-193, each incorporated herein by reference in their entirety. In turbulent reacting flows, the use of optical measurement techniques is therefore essential for reliable information. Laser-based tools are the method of choice offering the potential to measure most of the important quantities with high temporal and spatial resolution, often as one- or two-dimensional images, and the ability to perform the simultaneous detection of several quantities. See Eckbreth A. C., Laser Diagnostic for Combustion Temperature and Species, Gordon & Breach, 1996; Kohse-Höinghaus K., Jeffries J., Applied Combustion Diagnostics, Taylor & Francis, New York, 2002; Masri A. R., Dibble R. W, Barlow R. S., The structure of turbulent nonpremixed flames revealed by Raman-Rayleigh-LIF measurements, Prog. Energy Combust. Sci. 1996, 22, 307-362; and Wolfrum J., lasers in combustion: from basic theory to practical devices, Proc. Combust. Inst. 1998, 28, 1-41, each incorporated herein by reference in their entirety.
In recent years a variety of laser-based investigations in GT model combustors have been reported that, besides feasibility studies, concentrated on certain aspects of the combustion process or model validation. For example, Kaaling et al. performed temperature measurements with CARS in a RQL (rich-quench-lean) combustor, and Kampmann et al. used CARS simultaneously with 2-D Rayleigh scattering to characterize the temperature distribution in a double-cone burner. See Kaaling H., Ryden R., Bouchie Y., Ansart D., Magre P., Guin C., in: 13th International Symposium on Air Breathing Engines (ISABE), Chattanooga, Tenn. (USA), 1997 and Kampmann S., Seeger T., Leipertz A., simultaneous CARS and 2D laser ragleigh thermometry in a contained swirl combustor, Appl. Opt. 1995, 34, 2780-2786, each incorporated herein by reference. In the same combustor, Dinkelacker et al. studied the flame front structures and flame lift. See Dinkelacker F., Soika A., Most D., Hofmann D., Leipertz A., Polifke W., Döbbeling K., structure of locally quenched highly turbulent lean premixed flames, Proc. Combust. Inst. 1998, 27, 857-865, incorporated herein by reference in its entirety. Their experiments have been conducted at bluff-body-stabilized premixed methane/air flames, where flow and flame parameters have been varied systematically over a broad range of exit velocities and stoichiometries. They found that for this burner configuration not only one but two different liftoff criteria must be met. For very lean mixtures the chemically dominated ignition delay is found to be the rate-determining step. For other cases, the liftoff height can be determined by a newly described turbulent mixing dominated model. Fink et al. investigated the influence of pressure on the combustion process by applying PLIF of OH and NO in a LPP (lean pre-evaporized premixed) model combustor. See Fink R., Hupfer A., Rist D., in: Proceedings, ASME Turbo Expo, 2002, GT-2002-30078, incorporated herein by reference in its entirety. With respect to NOx reduction strategies, Cooper and Laurendeau performed quantitative NO LIF measurements in a lean direct injection spray flame at elevated pressures. See Cooper C. S., Laurendeau N. M., Comparison of laser-induced and planar laser-induced fluorescence measurements of nitric oxide in a high-pressure, swirl-stabilized, spray flame. Appl. Phys. 2000, B 70, 903-910 and Cooper C. S., Laurendeau N. M., Quantitative measurements of nitric oxide in high-pressure (2-5 atm), swirl-stabilized spray flames via laser-induced fluorescence, Combust. Flame. 2000, 123, 175-188, each incorporated herein by reference in their entirety. They have performed excitation scans and calibration comparisons to assess the background contribution for PLIF detection. Also, they presented and analyzed quantitative radial NO profiles measured by LIF so as to correct the PLIF measurements to within the accuracy bars of the LIF measurements via a single-point scaling of the PLIF image. Shih et al. applied PLIF of OH and seeded acetone in a lean premixed GT model combustor, and Deguchi et al. used PLIF of OH and NO in a large practical GT combustor. See Shih W. P., Lee J. G., Santavicca D. A., stability and emissions characteristics of a lean premixed gas turbine combustor, Proc. Combust. Inst. 1996, 26, 2771-2778 and Deguchi Y., Noda M., Fukuda Y., Ichinose Y., Endo Y., Inida M., Abe Y., Iwasaki S., industrial applications of temperature and species concentration monitoring using laser diagnostic, Meas. Sci. Technol. 2002, 13, R103-R115, each incorporated herein by reference in their entirety. They found that the automated LIBS unit is capable of monitoring trace element concentration fluctuations at ppb levels with a 1 min detection time under actual plant conditions. In addition, real-time measurement of O2 and CO concentrations in a commercial incinerator furnace was performed using TDLAS to improve the combustion control. The multiple-point laser measurement results to control secondary air allocation, higher secondary combustion efficiency was achieved, and CO concentration was reduced. Hedman and Warren used PLIF of OH, CARS, and LDV for the characterization of a GT-like combustor fired with propane in order to achieve a better understanding of the fundamentals of GT combustion. See Hedman P. O., Warren D. L., Turbulent velocity and temperature measurements from a gas-fueled technology combustor with a practical fuel injector, Combust. Flame. 1995, 100, 185-192, incorporated herein by reference in its entirety. PLIF of OH was also applied by Lee et al. to study flame structures and instabilities in a lean premixed GT combustor, by Arnold et al. to visualize flame fronts in a GT combustor flame of 400 kW, and by Fritz et al. for revealing details of flashback. See Lee S. Y., Seo S., Broda J. C, Pal S., Santoro R. J., An experimental estimation of mean reaction rate and flame structure during combustion instability in a lean premixed gas turbine combustor, Proc. Combust. Inst. 2000, 28, 775-782; Arnold A., Bombach R., Hubschmid W., Käppeli B., ERCOFTAC Bull, 1998, 38, 10-19; and Fritz J., Kröner M., Sattelmayer, in: Proceedings, ASME Turbo Expo 2001, 2001-GT-0054, each incorporated herein by reference in their entirety. Lofstrom et al. performed a feasibility study of two-photon LIF of CO and 2-D temperature mapping by LIF of seeded indium in a low-emission GT combustor. See Löfström C., Engström J., Richter M., Kaminsky C. F., Johansson P., Nyholm K., Nygren J., Aldén M., in: Proceedings, ASME Turbo Expo 2000, 2000-GT-0 124, incorporated herein by reference in its entirety. Four different laser diagnostic techniques were investigated in their work. The two more mature techniques, Planar Mie Scattering/Laser Induced Fluorescence and Planar Laser Induced Fluorescence of OH were used for fuel- and OH-visualisation, respectively. In addition, the applicability of some novel techniques in harsh industrial environments were investigated, two-line atomic fluorescence (TLAF) to obtain 2-dimensional temperature distributions, and two-photon LIF for the detection of CO. A comparison of two different laser excitation schemes for major species concentration measurements with laser Raman scattering was performed by Gittins et al. in a GT combustion simulator. See Gittins C. M., Shenoy S. U, Aldag H. R., Pacheco D. P., Miller M. F., Allen M. G., in: 38th AIAA Aerospace Sciences Meeting, Reno, Nev., 2000, incorporated herein by reference in its entirety. At a high-pressure test rig of the DLR, various laser techniques (LDV, CARS, PLIF of OH and kerosene, and 2-D temperature imaging via OH PLIF) were applied to GT combustors under technical operating conditions to achieve a better understanding of combustor behavior and to validate CFD codes. See Meier U. E., Wolff-Gaßmann D., Heinze J., Frodermann M., Magnusson I., Josefsson G., in: 18th International Congress on Instrumentation in Aerospace Simulation Facilities (ICIASF 99), Toulouse, 1999, pp. 7.1-7.7; Meier U. E., Wolff-Gaßmann D., Stricker W., LIF imaging and 2D temperature mapping in a model combustor at elevated pressure, Aerospace Sci. Technol. 2000, 4, 403-414; Carl M., Behrendt T., Fleing C., Frodermann M., Heinze J., Hassa C., Meier U., Wolff-Gaßmann D., Hohmann S, Zarzalis N., Gas Turbines Power, ASME J. Eng. 2001, 123, 810-816; and Kunz O., Noll B., Lückerath R., Aigner M., Hohmann S., in: 37th AIAA/ASME/SAE/ASEE Joint Propulsion Conference and Exhibition, Salt Lake City, Utah, 2001, AIAA 2001-3706, each incorporated herein by reference in their entirety.
Williams et al investigated syngas and methane flames for premixed swirl stabilized conditions for two different oxidizers of air and O2/CO2/N2. See Williams T. C., Shaddix C. R., Schefer R. W., Effect of Syngas Composition and CO2-Diluted Oxygen on Performance of a Premixed SwirlStabilized Combustor, Combust. Sci. Technol. 2008, 180, 64-88, incorporated herein by reference in its entirety. Simple flame images for different conditions have been presented along with exhaust gas emissions. They report lower nitrogen oxides concentrations (NOx) for the quasi-oxyfuel flames and higher carbon monoxide concentrations (CO), suggesting stoichiometric operation at 20-24% O2 as ideal for low emissions. Sautet et al. studied the length of natural gas/oxygen diffusion flames in a jet burner for free and confined configurations. See Sautet J. C., Salentey L., Ditaranto M., Samaniego J. M., length of natural gas-oxygen non-premixed flames, Combust. Sci. Technol. 2001, 166, 131-150, incorporated herein by reference in its entirety. Fuel jet Reynolds numbers were varied from 8362 to 16300 for five flames of which two were buoyancy controlled. The flame lengths were calculated from OH-chemiluminescence and indicated flames to be 2-3 times shorter than air flames with adiabatic flame temperatures in the region of 3050 K. Ditaranto and Hals discussed the effect of stoichiometric operation and high O2 content in oxidizer on thermo-acoustic oscillations in sudden expansion jet configuration. See Ditaranto M., Hals J., Combustion instabilities in sudden expansion oxy-fuel flames. Combust, Flame. 2006, 146, 493-512. dx.doi.org/10.1021/ef300539c|Energy Fuels 2012, 26, 4599-4606, incorporated herein by reference in its entirety. They reported occurrence of thermo-acoustic instabilities as O2 content in the oxidizer was increased, characterizing different instability modes dependant on flow velocity and flame speed variations. The discussion above focused on oxy-fuel combustion using ion transport membranes in order to separate oxygen from air, then using the oxygen that permeates through the membrane in a combustion process in the other side of the membrane as shown in FIG. 6 FIG. 6 is an illustrative flowsheet for oxy-fuel combustion process using membrane reactor technology, with additional unit operations for carbon capture. See Habib M. A., Badr H. M., Ahmed S. F., Ben-Mansour R., Mezghani K., Imashuku S., lao G. J., Shao-Horn Y., Mancini N. D., Mitsos A., Kirchen P., Ghoniem A. F., A review of recent developments in carbon capture utilizing oxy-fuel combustion in conventional and ion transport membrane systems, Int. J. Energy Res. 2011, 35, 741-764, incorporated herein by reference in its entirety.
More recently, strong demand for tonnage quantities of oxygen is encouraged by the steady growth in chemical process operations. For instance, oxyfuel combustion process and oxygen-blown gasification to convert coal and natural gas into an intermediate synthesis gas that can be further processed to produce electricity, chemicals and transportation fuels. There have been two fundamental approaches to air separation, which are cryogenic and non-cryogenic distillation. The cryogenic distillation is typically reserved for applications that require tonnage quantity of oxygen at ultra-low temperature. The latter involves the separation of air at ambient temperatures using either molecular sieve adsorbents via pressure swing adsorption (PSA), or membrane separation process using the polymeric membranes. Recently, a third category of air separation has emerged, which is based on specialized ceramic membranes that separate oxygen from air at elevated temperatures, in contrast to the super-cooled temperatures required by conventional cryogenic distillation. This novel technique is based on dense ceramic membranes, which carry out the separation of oxygen from air at elevated temperatures, typically 800 to 900° C. MIEC (Mixed Ionic Electronic Conducting) membranes, ITM (Ion Transport Membranes), and OTM (Oxygen Transport Membranes) are acronyms that used to refer to high temperature ceramic membranes. See Hashim S. M., Mohamed A., Bhatia S., Current status of ceramic-based membranes for oxygen separation from air, Advances in Colloid and Interface Science. 2010, 160, 88-100, incorporated herein by reference in its entirety. These terms will be used throughout this work.
Ceramic based membranes for oxygen separation systems can be categorized into pure oxygen conducting membranes and mixed ionic-electronic conducting membranes. The solid electrolytes are pure oxygen conducting membranes, where electrodes are provided for the electron pathway. See Sunarso J., Baumann S., Serra J. M., Meulenberg W. A., Liu S., Lin Y. S., Diniz da Costa J. C., Mixed ionic-electronic conducting (MIEC) ceramic-based membranes for oxygen separation, J Membr Sci. 2008, 320:13-41, incorporated herein by reference in its entirety. The main advantage of this system is the control over the amount of oxygen generated via the application of an electric current. Compared to solid electrolytes, mixed ionic-electronic conducting membranes require neither electrodes nor an external circuit to operate. The electronic conductivity itself performs as an internal short circuit involving oxygen partial pressure gradient. Oxygen ions permeate from the high oxygen partial pressure side to the low oxygen partial pressure side, whilst the overall charge neutrality is maintained by a counterbalancing flux of electrons, as idealized schematically in FIG. 7. FIG. 7 is a schematic diagram of a dense ceramic membrane based on conduction mechanism. It should be noted that oxygen separation through this process has the advantage of producing high-purity oxygen.
It is worth noticing that ceramic materials with mixed ionic-electronic conducting characteristics typically have defined phase structures that can be derived from perovskite, fluorite, brownmillerite, and other similar types of materials. See Shao Z. P., Yang W. S., Cong Y., Dong H., Tong J. H., Xiong G. X., Investigation of the permeation behavior and stability of a Ba0.5Sr0.5Co0.8Fe0.2O3−δ oxygen membrane, J Membr Sci. 2000, 172, 177-188; Kharton V. V., Viskup A. P., Kovalevsky A. V., Naumovic E. N., Marques F. M. B., Ionic transport in oxygen-hyperstoichiometric phases with K2NiF4-type structure, Solid State Ionics. 2001, 143, 337-353; Ting C., Hailei Z. Nansheng X., Yuan L., Xionggang L. U., ˜Weizhong D., ˜Fushen L., Synthesis and oxygen permeation properties of a Ce0.8Sm0.2O3−δ LaBaCO2O5+δ, Journal of Membrane Science 2011, 370, 158-165; and Wiik K., Aasland S., Hansen H. L., Tangen L. L., Odegard R., Oxygen permeation in the system SrFeO3−x—SrCoO3−y, Solid State Ionics. 2002, 152-153, 675-680, each incorporated herein by reference in their entirety. Among ceramic membranes with mixed ionic-electronic conducting characteristics, perovskite-type and fluorite-type are the best structures in case of oxygen permeation properties; however, the perovskite type ceramic membranes have higher permeability and promising potential for improvement. See Fan C. G., Zuo Y. B., Li J. T., Lu J. Q., Chen C. S., Bae D. S., Highly permeable La0.2Bao0.8Co0.8Fe0.2−xZrxO3−δ membranes for oxygen separation, Sep. Purif Technol. 2007, 55, 35; Ishihara T., Yamada T., Arikawa H., Nishiguchi H., Takita Y., mixed electronic-oxide ionic conductivity and oxygen permeating property of Fe-, Co- or Ni-doped LaGaO3 provskite oxide, Solid State Ionics. 2000, 135, 631-636; and Fan C. G., Zuo Y. B., Li J. T., Lu J. Q., Chen C. S., Bae D. S., Highly permeable La0.2Ba0.8Co0.8Fe0.2−xZrxO3−δ membranes for oxygen separation, Sep Purif Technol. 2007, 55, 35, each incorporated herein by reference in their entirety.
Based on the difference in the oxygen chemical potentials between the feed side and the permeate side, the membrane temperature and its ambipolar conductivity, oxygen migrates from the high pressure feed side to the low pressure permeate side, according to the overall transport processes summarized as follows:
i) gaseous oxygen mass transfer (advection and diffusion) from the feed stream to the membrane surface, adsorption onto the membrane surface, dissociation and ionization of oxygen molecules and subsequent incorporation of the ions into the lattice vacancies (feed side surface exchange),
ii) Transport of lattice oxygen ions through the membrane (bulk diffusion),
iii) Association of lattice oxygen ions to oxygen molecules and desorption from the membrane surface into the gas phase (permeate side surface exchange), gaseous oxygen mass transfer (advection and diffusion) from the membrane surface to the permeate stream. See Manning P. S., Sirman J. D., Kilner J. A., Oxygen self-diffusion and surface exchange studies of oxide electrolytes having the fluorite structure, Solid State Ionics. 1996, 93(1-2), 125-132; Ishihara T., Kilner J. A., Honda M., Sakai N., Harumi Y., Yusaku T., Oxygen surface exchange and diffusion in LaGaO3 based perovskite type oxides, Solid State Ionics. 1998, 113-115, 593-600; Ruiz-Trejo E., Sirman J. D., Baikov Y. M., Kilner J. A., Oxygen ion diffusivity surface exchange and ionic conductivity in single crystal Gadolinia doped Ceria, Solid State Ionics. 1998, 113-115, 565-569; and Lane J. A, Kilner J. A., Oxygen surface exchange on gadolinia doped ceria, Solid State Ionics. 2000, 136-137, 927-932, each incorporated herein by reference in their entirety. FIG. 8 is a schematic diagram of oxygen permeation through mixed ionic-conducting membrane. FIG. 8 demonstrates that oxygen permeation through a dense mixed ionic-electronic conducting material is limited by surface exchange resistance, bulk diffusion limitations, or both. See Tan X., Liu Y., Li K., Mixed conducting ceramic hollow fibre membranes for air separation, AIChE J. 71 (2005) 1991; Kim S., Yang Y. L., Jacobson A. J., Abeles B., Diffusion and surface exchange coefficients in mixed ionic electronic conducting oxides from the pressure dependence of oxygen permeation, Solid State Ionics. 1998, 106:189-195; and Lin Y. S., Wang Y., Han J., Oxygen permeation through thin mixed-conducting solid oxide membranes. AIChE J. 1994, 40:786-798, each incorporated herein by reference in their entirety. It should be noted that the bulk diffusion will be the controlling step when the membrane is relatively thick.
FIG. 8 visualizes the mechanism for oxygen permeation through a mixed ionic-electronic conducting membrane. It can be seen that the permeation process from the high oxygen partial pressure side to the low oxygen partial pressure side by dividing the process into three zones: (1) an interfacial zone on the high partial pressure or air side; (2) a Central bulk zone; and (3) an interfacial zone on the low oxygen partial pressure or sweep gas side. The demonstration of incorporating both bulk diffusion and surface exchange kinetics into a single unambiguous equation has been done by few research groups. For example, the following general assumptions were used for the derivation of Eq. (2.1) by Tan and Li in the formation of mathematical models for the perovskite systems: 1) the oxygen permeation flux is controlled by the surface exchange reactions. 2) The operation is under steady-state isothermal operation. 3) The radial diffusion of gases is neglected. 4) Ideal gas law is applied to the gas phase. 5) The mass transfer resistance of gas phase to oxygen permeation is negligible and the oxygen partial pressures on both shell side and tube side of the membrane surfaces are the same.
                                          dNo            2                    dl                =                                            K              f                        ⁡                          [                                                                    (                                                                  P                        ′                                            ⁢                                              o                        2                                                              )                                    0.5                                -                                                      (                                                                  P                        ″                                            ⁢                                              o                        2                                                              )                                    0.5                                            ]                                                                                            K                  f                                ⁢                                  ln                  ⁡                                      (                                                                  r                        2                                            /                                              r                        1                                                              )                                                  ⁢                                                      (                                                                  P                        ′                                            ⁢                                              o                        2                                                              )                                    0.5                                ⁢                                                      (                                                                  P                        ″                                            ⁢                                              o                        2                                                              )                                    0.5                                                            π                ⁢                                                                  ⁢                                  D                  v                                                      +                                                            (                                                            P                      ′                                        ⁢                                          o                      2                                                        )                                0.5                                            2                ⁢                π                ⁢                                                                  ⁢                                  r                  1                                                      +                                                            (                                                            P                      ″                                        ⁢                                          o                      2                                                        )                                0.5                                            2                ⁢                π                ⁢                                                                  ⁢                                  r                  2                                                                                        (        2.1        )            where N is the molar flow rate, 1 is the variable length of hollow fiber membrane, kf is the forward reaction rate constant, Dv is the oxygen vacancy diffusion coefficient, p′o2 and p˜o2 are partial pressures of oxygen at the feed and permeate side, respectively and r2 and r1 are the outer and inner diameter radius of the membrane tube.
Eq. (2.1) is specially derived for tubular membranes and has been applied successfully in the hollow-fiber systems of Ba0.5Sr0.5Co0.8Fe0.2O3−δ (BSCF) and La0.6Sr0.4Co0.2Fe0.8O3−δ—(LSCF). See Xu S. J., Thomson W. J., Oxygen permeation rates through ion-conducting perovskite membranes, Chem Eng Sci. 1999, 54: 3839 and Lee T., Yang Y., Jacobson A., Abelesa B., Zhou M., Oxygen permeation in dense SrCo0.8Feo2O3−δmembranes: Surface exchange kinetics versus bulk diffusion, Solid State Ionics. 1997, 100, 77-85, each incorporated herein by reference in their entirety. The detailed derivation of Eq. (2.1) can be found elsewhere. See Shao Z., Xiong G., Tong J., Dong H., Yang W., ˜Ba effect in doped Sr(Co0.8Fe0.2)O3−δ on the phase structure and oxygen permeation properties of the dense ceramic membranes', Sep. Purif. Technol., 25, 419-429 (2001); Wang H., Wang R., Liang D., Yang W., Experimental and modeling studies on Ba0.5Sr0.5Co0.8Fe0.2O3−δ (BSCF) tubular membranes for air separation, J. membr Sci. 2004, 243, 405; and Ge L., Shao Z., Zhang K., Ran R., Diniz da Costa J., Liu S., Evaluation of mixed-conducting lanthanum-strontium-cobaltite ceramic membrane for oxygen separation, AIChE J. 2009, 55, 2603, each incorporated herein by reference in their entirety. The use of Eq. (2.1) for the perovskite systems is very useful in the scaled-up engineering calculations, wherein the oxygen permeation flux performance inside the hollow-fiber perovskite membrane modules, which consists of many small-long tubes, can be estimated. Consequently, the oxygen permeation flux with respect to log mean membrane area, dAm=2πrmdl where rm=(r2−r1)/ln(r2/r1) can be expressed by:
                    ⁢          (      2.2      )                  Jo      2        =                            K          f                ⁢                              D            v                    ⁡                      (                                                            (                                                            P                      ′                                        ⁢                                          o                      2                                                        )                                0.5                            -                                                (                                                            P                      ″                                        ⁢                                          o                      2                                                        )                                0.5                                      )                                                2          ⁢                      (                                          r                2                            -                              r                1                                      )                    ⁢                                                    K                f                            ⁡                              (                                                      P                    ′                                    ⁢                                      o                    2                                    ⁢                                      P                    ″                                    ⁢                                      o                    2                                                  )                                      0.5                          +                                            r              m                                      r              1                                ⁢                                                    D                v                            ⁡                              (                                                      P                    ′                                    ⁢                                      o                    2                                                  )                                      0.5                          +                                            r              m                                      r              1                                ⁢                                                    D                v                            ⁡                              (                                                      P                    ″                                    ⁢                                      o                    2                                                  )                                      0.5                              
For the sake of analyses, it is assumed that the oxygen permeation in Eq. (2.2) is still applicable to other mixed conducting membranes such as SrCo0.9Nb0.1O3−δ and La0.2Ba0.8Co0.8Fe0.2−xZrxO3−δ. See Ito W., Nagai T., Sakon T., Oxygen separation from compressed air using a mixed conducting perovskite-type oxide membrane, Solid State Ionics. 2007, 178, 809, incorporated herein by reference in its entirety. Lc is defined by the membrane thickness at which the oxygen permeation resistance by bulk diffusion equals by the surface exchange reactions. Lc is expressed as:
                              L          c                =                                            D              v                                      2              ⁢                                                          ⁢                              K                f                                              ⁢                      (                                          1                                                      P                    ′                                    ⁢                                      o                    2                    0.5                                                              +                              1                                                      P                    ″                                    ⁢                                      o                    2                    0.5                                                                        )                                              (        2.3        )            
It is noted in Eq. (2.3) that value depends on oxygen partial pressure and temperature. A larger characteristic thickness is resulted from higher and operating at lower temperature or lower oxygen partial pressure. When the membrane has a thickness far less than the critical thickness, Lc, the resistance by bulk diffusion can be neglected; therefore, the surface exchange reaction becomes the rate-limiting step. Lee et al. indicated that the permeation in membranes in the thickness range of 1 to 2.6 mm was controlled both by the bulk diffusion of oxide ions and by surface exchange. In this case, Eq. (2.1) can be further simplified as:
                                          dNo            2                    dl                =                                            K              f                        ⁢                          ⌊                                                                    (                                                                  P                        ′                                            ⁢                                              o                        2                                                              )                                    0.5                                -                                                      (                                                                  P                        ″                                            ⁢                                              o                        2                                                              )                                    0.5                                            ⌋                                                                                            (                                                            P                      ′                                        ⁢                                          o                      2                                                        )                                0.5                                            2                ⁢                π                ⁢                                                                  ⁢                                  r                  1                                                      +                                                            (                                                            P                      ″                                        ⁢                                          o                      2                                                        )                                0.5                                            2                ⁢                π                ⁢                                                                  ⁢                                  r                  2                                                                                        (        2.4        )            See Zhu X., Sun S., Cong Y., Yang W., Operation of perovskite membrane under vacuum and elevated pressures for high-purity oxygen production, J Membr Sci. 2009, 345, 47-52, incorporated herein by reference in its entirety.
Surface exchange reactions may involve many sub-steps, which are oxygen transfer from the gas phase to membrane phase, physical adsorption on surface, dissociation with electronic transfer to yield chemisorbed oxygen species, and the incorporation in surface layer including the reverse reactions. See Zhu X., Cong Y., Yang W., Oxygen permeability and structural stability of BaCe0.15Fe0.85O3−δ membranes, J Mem br Sci. 2006, 283, 38-44, incorporated herein by reference in its entirety.
The oxygen permeation flux is inversely proportional to the membrane thickness; thus, reducing the membrane thickness will increase the oxygen permeation as long as bulk diffusion prevails. Kim et al. derived an equation applicable in tubular perovskite membranes to explain the oxygen permeation data in Sr0.5Co0.8Fe0.2O3−δ and Sm0.5 Sr0.5CoO3−δ in the surface exchange reaction limited. Eq. (2.5) can be used for surface exchange limited region in tubular perovskite membranes.
                              Jo          2                =                                            π              ⁢                                                          ⁢                              r                1                            ⁢                              r                2                            ⁢              k                                                      (                                                      r                    1                                    +                                      r                    2                                                  )                            ⁢              S                                ⁢                      (                                                                                                      P                      ′                                        ⁢                                          o                      2                                                                                                                                  P                      o                                        ⁢                                          o                      2                                                                                  -                                                                                          P                      ″                                        ⁢                                          o                      2                                                                                                                                  P                      o                                        ⁢                                          o                      2                                                                                            )                                              (        2.5        )            where, k is the surface exchange coefficient unit and S is the effective membrane area. When a membrane is relatively thick, the overall oxygen permeation is controlled by the bulk diffusion. The flux in this case is generally described by Wagner theory. The flux of oxygen, Jo2 is given by Eq. (2.6):
                              Jo          2                =                              1                                          4                2                            ⁢                              F                2                            ⁢              L                                ⁢                                    ∫                                                μ                                      o                    2                                                  ⁡                                  (                  2                  )                                                            μ                                                      o                    2                                    ⁡                                      (                    1                    )                                                                        ⁢                                          t                i                            ⁢                              t                e                            ⁢                              σ                t                            ⁢              d              ⁢                                                          ⁢                              μ                                  o                  2                                                                                        (        2.6        )            where F (C/mol) the Faraday constant; L (m) the membrane thickness and t is the product of the electronic transference number te, ionic transference number ti and total conductivity σt. Lin et al. and Qi et al. derived the oxygen permeation flux equations for ionic or mixed-conducting ceramic membranes in terms of electrical conductivity. See Lin Y. S., Wang Y., Han J., Oxygen permeation through thin mixed-conducting solid oxide membranes, AIChE J. 1994, 40:786-798 and Qi X., Lin Y., Swartz S., electrical transport and oxygen permeation properties of lanthanum cobaltite membranes synthesized by different methods, Ind Eng Chem Res. 2000, 39, 646, each incorporated herein by reference in their entirety. In this case, the oxygen permeation flux can be expressed as:
                              Jo          2                =                              RT                          4              ⁢                                                          ⁢                              F                2                            ⁢              L                                ⁡                      [                                                                                σ                    h                                    ⁡                                      (                                                                  P                        ′                                            ⁢                                              o                        2                                                              )                                                  ⁢                                  (                                      1                    -                                                                  (                                                                                                            P                              ″                                                        ⁢                                                          o                              2                                                                                                                                          P                              ′                                                        ⁢                                                          o                              2                                                                                                      )                                                                    1                        /                        4                                                                              )                                            +                                                                    σ                    e                                    ⁡                                      (                                                                  P                        ′                                            ⁢                                              o                        2                                                              )                                                  ⁢                                  (                                                                                    (                                                                                                            P                              ″                                                        ⁢                                                          o                              2                                                                                                                                          P                              ′                                                        ⁢                                                          o                              2                                                                                                      )                                                                    1                        /                        4                                                              -                    1                                    )                                                      ]                                              (        2.7        )            where, R (J/mol K) is the gas constant and T (K) is the temperature. Eq. (2.7) has been used successfully by Rui et al. to describe the oxygen permeation through dense ceramic membranes with finite rate of chemical reaction. See Rui Z., Li Y., Lin Y., Analysis of Oxygen Permeation through Dense Ceramic Membranes with Chemical Reactions of Finite Rate, Chem Eng Sci. 2009, 64, 172, incorporated herein by reference in its entirety. Oxygen permeation flux within Bi1.5Y0.3Sm0.3O3−δ (t=1.2 mm, T=873-953 K) and La0.6Sr0.4Co0.2Fe0.8O3−δ (t=1.12 mm, T=1273 K) is well explained by this equation. Although Eq. (2.7) is originally an empirical equation, it has been applied in the literature with reasonable predictions. See Akin F. T., Jerry, Lin Y. S., Oxygen permeation through oxygen ionic or mixed-conducting ceramic membranes with chemical reactions, Journal of Membrane Science. 2004, 231, 133-146, incorporated herein by reference in its entirety.
Wang et al. revealed that the controlling step of the oxygen permeation for the 1.5 mm thickness BSCF tubular membrane was bulk diffusion at the temperature range of 700° C. to 900° C.; hence, the surface exchange reaction does not favor the oxygen permeation flux. The oxygen permeation flux is described in Eq. (2.8).
                              Jo          2                =                                            π              ⁢                                                          ⁢                              LC                i                            ⁢                              D                a                                                    2              ⁢                                                          ⁢              S              ⁢                                                          ⁢                              ln                ⁡                                  (                                                            r                      1                                        /                                          r                      2                                                        )                                                              ⁢                      ln            ⁡                          (                                                                    P                    ″                                    ⁢                                      o                    2                                                                                        P                    ′                                    ⁢                                      o                    2                                                              )                                                          (        2.8        )            where Ci stands for density of oxygen ions and Da for ambipolar oxygen ion-hole diffusion coefficient, Da could be determined from oxygen permeation flux data. Ge et al. theoretically analyzed that the oxygen permeation flux through La0.4Sr0.6CoO3−δ disk-shaped membrane at 950° C. was controlled by both the bulk diffusion and oxygen surface exchange reactions; therefore, the oxygen permeation flux is given as:
                              Jo          2                =                                            D              v                                      2              ⁢                                                          ⁢              L                                ⁢                                                    K                f                                                              1                                                            (                                                                        P                          ′                                                ⁢                                                  o                          2                                                                    )                                        0.5                                                  +                                  1                                                            (                                                                        P                          ″                                                ⁢                                                  o                          2                                                                    )                                        0.5                                                                                                      K                f                                                              1                                                            (                                                                        P                          ′                                                ⁢                                                  o                          2                                                                    )                                        0.5                                                  +                                  1                                                            (                                                                        P                          ″                                                ⁢                                                  o                          2                                                                    )                                        0.5                                                  +                                                      D                    v                                                        2                    ⁢                                                                                  ⁢                    L                                                                                ⁢                      (                                          C                v                                  ′                  ⁢                                                                          ⁢                  e                                            -                              C                v                e                                      )                                              (        2.9        )            
Where Ce and C′e are the oxygen vacancy concentrations in the material under thermal equilibrium with the atmosphere surrounded by oxygen partial pressure at feed side and oxygen partial pressure at the permeate side and (P′o2 and P″o2), respectively. Dv and kf can be determined by proper fitting of the experimental data into Eq. (2.9), which requires prior knowledge of (a) and values and (b) the oxygen permeation fluxes under the applied oxygen partial pressure gradient across the membrane at selected temperatures by oxygen permeation experiments.
Chang et al. systematically compared the performance of symmetric (1.5 mm thick) and asymmetric mixed-conducting membranes (200 μm-thick thin dense layer and 1.3 mm-thick support) with correlation of overall oxygen permeation resistance across the membrane. See Chang X., Zhang C., He Y., Dong X., Jin W., Xu N., A Comparative Study of the Performance of Symmetric and Asymmetric Mixed-conducting Membranes, Chin J Chem Eng. 2009, 17, 562, incorporated herein by reference in its entirety. They prepared the asymmetric membrane consists of the support and thin dense layer from the same composition SrCo0.4 Fe0.5Zr0.1O3−δ (SCFZ) perovskite-type oxide. The oxygen permeation flux in both membrane architectures is represented as:
                              Jo          2                =                              1            S                    ⁢                      RT                                          R                p                            +                              R                d                                              ⁢                      ln            ⁡                          (                                                                    P                    ′                                    ⁢                                      o                    2                                                                                        P                    ″                                    ⁢                                      o                    2                                                              )                                                          (        2.10        )            
Where Rp and Rd are the resistance in the porous support and thin dense layer, respectively. For an asymmetric membrane, the overall resistance is the sum of the resistance in the porous support and in the thin dense layer (Roverall=Rp+Rd). In the symmetric membrane Roverall=Rp. They concluded that the oxygen permeation flux on the asymmetric membrane was higher than the symmetric membrane due to the significant decrease of bulk diffusion resistance in the thin dense layer of the asymmetric membrane. Chang et al. reported detailed calculation of the overall oxygen permeation resistance and modeling of the transport resistance through a membrane.
The performance of oxygen permeation flux depends on both membrane materials and the membrane thickness. The permeation conditions such as pressure and temperature also influence the flux according to the Eq. (2.7); in addition, the sweep gas flow rate also plays an important role. The dependence of oxygen permeation flux on the thickness of the disk membrane of 0.55, 1.10 and 1.65 mm has been observed by Sunarso et al. See Sunarso J., Liu S., Lin Y., Diniz da Costa J., Oxygen permeation performance of BaBiO3-δ ceramic membranes, J Membr Sci. 2009, 344, 281-287, incorporated herein by reference in its entirety. FIG. 9 is a graph of the oxygen permeation fluxes as function of temperature at different membrane thicknesses. As can be seen in FIG. 9, the oxygen permeation flux was insignificant when the temperature was lower than 600° C.; however, at temperature higher than 650° C., considerable oxygen permeation flux value was obtained. These researchers claimed that the sharp increase of oxygen permeation flux was obtained when the temperature was above 800° C. due to the improvement of oxygen ionic diffusion and/or surface reaction rate. It can be obviously seen the six fold enhancement (from 0.2 to 1.2 ml/cm2min) for 0.55 mm disk membrane thickness when the temperature increased from 800 to 950° C. Sunarso et al. also revealed that the oxygen permeation flux increased as the membrane thickness decreased, especially when the temperature was higher than 800° C. Interestingly, they observed that the increase of oxygen permeation flux was not inversely proportional to the membrane thickness reduction. This can be explained by reducing the thickness of the disk membrane by half, which was from 1.10 mm to 0.55 mm; as a result, the increase of oxygen permeation fluxes was only 20%, rather than 100%. From the results obtained, Sunarso et al. suggested that both bulk diffusion and surface exchange kinetics were the rate limiting steps for the oxygen transport across the membranes. The similar observation was reported by Watanabe et al. See Watanabe K., Takauchi D., Yussa M., Kida T., Shimanoe K., Teraoka Y., et al., Oxygen Permeation Properties of Co-Free Perovskite-Type Oxide Membranes Based on BaFe1−yZryO3−δ, J Electrochem Soc. 2009, 156, 81, incorporated herein by reference in its entirety.
Zhu et al. investigated the performance of oxygen permeation flux of dense BaCe0.15Fe0.8O3−δ (BCF1585) ceramic membranes synthesized by the solid-state reaction (SSR) method and EDTA-citric acid (EC) method. See Zhu X., Cong Y., Yang W. J., Effects of synthesis methods on oxygen permeability of BaCe0.15Fe0.85O3−δ ceramic membranes, Mem br Sci. 2006, 283, 158-163, incorporated herein by reference in its entirety. FIG. 10 is a graph of the dependence of oxygen permeation flux of the membranes synthesized by different methods on temperature. FIG. 10 shows the dependence of oxygen permeation flux of BCF 1585 membranes synthesized by different methods on the operating temperatures. FIG. 10 shows the oxygen permeation flux of the membrane synthesized by SSR method was higher than the membrane synthesized by EC method. They observed the oxygen permeation flux for the SSR method derived membrane was 0.92 ml/cm2·min compared to EC method derived membrane with 0.71 ml/cm2·min at 940° C., respectively. It has been reported by Han et al. that oxygen partial pressure in the permeate side could reach as low as 10−13 atm. under the vacuum operation model; as a result, the oxygen permeation flux would be extremely high. See Han J.,; Xomeritakis G., Lin Y. S., Oxygen permeation through thin zirconia/yttria membranes prepared by EVD, Solid State Ionics. 1997, 93, 263-272, incorporated herein by reference in its entirety. It is vitally important to note that the mechanical strength of thinner perovskite membranes under such condition is very poor; therefore, the deposition of a thin membrane on a porous support or asymmetric structure seems to be a wise solution.
Li and Tan found that the LSCF hollow-fiber module operated at elevated pressure may not be a good choice compared to the operation under vacuum conditions. They observed that the oxygen flux of 0.31 ml/cm2·min can be obtained when the vacuum pressure in the tube side of the module is reduced to 0.01 atm. additionally; they claimed that it would not be possible to achieve such oxygen permeation flux at elevated pressure operation even the shell side pressure is increased to 30 atm. It can be noted that the elevated pressure operation is generally adopted in the conventional polymeric hollow-fiber membrane modules for gas separation. Therefore, the vacuum operation in the permeate side of the membrane is more preferable. In addition, Li and Tan claimed that the improvement in oxygen flux can be observed when the pressure in the permeate side is decreased, but only up to a certain pressure value or effective pressure. The enhancement in level improves the performance in oxygen permeation flux because the driving force across the membrane is increased. However, there is no improvement in oxygen permeation flux with further decrease in the vacuum.
Zydorczak et al. studied the performance of oxygen permeation flux in an ultra-thin La0.6Sr0.4Co0.2Fe0.8O3−δ. by investigating the effects of operating temperatures and sweep gas flow rate (helium) on the oxygen permeation fluxes. See Zydorczak B., Wu Z., Li K., Fabrication of ultrathin LSCF hollow fibre membranes for oxygen permeation, Chem Eng Sci. 2009, 64, 4383, incorporated herein by reference in its entirety. They observed that the oxygen permeation flux increased when the operating temperature was increased. The similar trend occurs with the increasing of sweep gas flow rates. The oxygen partial pressure difference across the hollow fiber membrane is increased when flow rate of sweep gas increased. The oxygen partial pressure gradient acts as the driving force for the transport of oxygen through the membrane; therefore, the sharp increase in oxygen permeation flux can be observed. The reduced membrane thickness does contribute to the improved oxygen permeation flux, than that reported in previous studies by Tan et al. and Wang et al., See Tan X., Wang Z., Liu H., Liu S., Enhancement of oxygen permeation through La0.6Sr0.4Co0.2Fe0.8O3−δ hollow fibre membranes by surface modifications, J Mem br Sci. 2009, 324, 128 and Wang Z., Yang N., Meng B., Tan X., Li K., Preparation and oxygen permeation properties in hollow fibre membranes, Ind Chem Eng Res. 2009, 48, 510, each incorporated herein by reference in their entirety. It is important to note that surface exchange reaction and bulk diffusion are two factors that contribute to the transport resistance with respect to the operating temperature. Surface exchange reaction is the rate limiting step at lower temperatures; conversely, bulk diffusion becomes the rate limiting step with the increase in the temperature. The similar observation has been reported by Xu and Thomson. The hollow fiber membrane developed is extremely thin; hence, at high temperature, the flux will increase rapidly. They concluded that the reduced thickness of the membrane has an enormous effect on oxygen permeation flux as the ionic bulk diffusion will be the rate limiting step at high temperature.
Another study done by Tan and Li revealed the effect of sweep gas operation to the performance of oxygen permeation flux in hollow fiber modules. See Tan X., Li K., Oxygen production using dense ceramic hollow fiber membrane modules with different operating modes, AIChE J. 2007, 53, 838, incorporated herein by reference in its entirety. They observed that the oxygen concentration in the permeate side decreased when the argon flow rate is increased for all the operating temperatures. They reported that the use of water vapor as a sweep gas in investigating the oxygen flux gave the same operating performance as under argon gas condition. Wang et al. investigated the production of high-purity oxygen by BCFZ perovskite hollow fiber membranes swept with steam; consequently, they found that the oxygen permeation obtained was 4.1 ml/cm2min at 875° C. See Wang H., K'olsch P., Schiestel T., Tablet C., Werth S., Caro J., Production of high-purity oxygen by perovskite hollow fiber membranes swept with steam, J Membr Sci. 2006, 284, 5-8, incorporated herein by reference in its entirety. Hence, they concluded that the steam employed as a sweep gas has a promising potential to produce high-purity oxygen with reasonable oxygen permeation flux due to the problems related to high temperature sealing can be solved. However, two fundamental problems of employing steam or water vapor are membrane stability and membrane module itself. Tan et al. and Wang et al. did not explain these two problems due to the limitation of experimental apparatus.
Mixed ionic and electronic conducting ceramic membranes have received increasing interest from academia and industry. A major industrial effort is currently devoted to the development of the mixed-conducting ceramic membrane reactor technology for combustion of hydrocarbons.
Permeation Equation:
In the ceramic membrane reactor, oxygen permeates through the ionic or mixed-conducting ceramic membrane via a complex mechanism. It usually includes adsorption of oxygen and charge transfer reaction on the membrane surface exposed to air, oxygen vacancy and electron diffusion in the membrane bulk, and charge transfer and chemical reaction on the membrane surface exposed to a reducing gas. The detailed mathematical formulation for oxygen permeation through mixed-conducting ceramic membranes is fairly complex. The oxygen permeation through oxygen ionic or mixed-conducting ceramic membranes under reaction conditions was also examined by Zebao et. al. with a model taking into account of different electrical transport mechanisms (p-type and n-type transports) and finite reaction rate. See Zebao R., Yongdan L., Lin Y. S., Analysis of oxygen permeation through dense ceramic membranes with chemical reactions of finite rate, Chemical Engineering Science. 2009, 64, 172-179, incorporated herein by reference in its entirety. It was demonstrated in their work that with a reaction consuming oxygen in one side of the membrane, the oxygen partial pressure in the reaction side decreases and the oxygen permeation flux increases with the increase in the reaction rate for both the p-type and the n-type transport dominated mechanism. Also they reported that the increase in reaction rate causes a transition of the transport mechanism from p-type to n-type. This transition leads to an increase in the permeation flux by up to 30 times. This effect offers one explanation for the large discrepancies in published permeation data for membrane reactors of partial oxidation reaction employing an oxygen permeable dense ceramic membrane. They reported also, for a membrane with a specific transport mechanism, the increase in the reactant partial pressure causes an increase in the reaction rate and oxygen permeation flux. However, the increase in the inlet inert gas amount has a complicated effect on the oxygen permeation flux because it lowers both oxygen partial pressure and the reaction rate at the same time.
The OCM (oxidative coupling of methane) and SOE (selective oxidation of ethane) reactions involve oxidative reactions of methane or ethane to form ethylene as the intermediate (desired) product. The final (equilibrium) product is carbon dioxide (and water). Both the OCM and SOE reaction mechanisms are very complex and may involve over hundreds of steps. See Akin F., Lin Y., oxidative coupling of methane in dense ceramic membrane reactor with high yields, AIChE J. 2002, 48, 2298-2306 and Akin F. T., Lin Y. S., Selective oxidation of ethane to ethylene in a dense tubular membrane reactor, J. Membrane Sci. 2002, 209, 457-467, each incorporated herein by reference in their entirety. Akin and Jerry demonstrated how the extent of the reaction (or reactivity) and reactant flow rate affect the oxygen permeation flux, they used the following simple reaction to exemplify the complex oxidative reactions of methane or ethane to ethylene and finally to carbon dioxide: See Akin F. T., Jerry, Lin Y. S., Oxygen permeation through oxygen ionic or mixed-conducting ceramic membranes with chemical reactions, Journal of Membrane Science. 2004, 231, 133-146, incorporated herein by reference in its entirety.CO+1/2O2→CO2  (A)
Where, carbon monoxide represents a hydrocarbon reactant. The use of this simple reaction can help in obtaining semi-analytical solution for the combustion with oxygen permeation. The basic principle illustrated with this simple reaction can be extended for the more complex reaction systems if detailed reaction mechanism and kinetic equations are known. In order to obtain analytical expressions for the oxygen partial pressure in the reaction chamber, Akin and Jerry used a simple reactor model, continuously stirred tank reactor (CSTR), to describe the reaction chamber in this work. Since in most laboratory studies, membrane reactor experiments were conducted on disk-shaped or short tubular membrane, such as the BYS used in their work, the CSTR model can catch the major characteristics of the oxidation reaction in the membrane reaction chamber. Instead of putting specific reaction kinetics in the model in account for reaction rate, they only considered two extreme cases in the modeling and analysis: (a) extremely fast reaction rate, i.e., complete conversion of the reactant (CO in this work) with oxygen permeating to the reaction chamber, (b) extremely slow reaction rate, i.e., no reaction between the oxygen permeating with the reactant (CO). In the second case, the reactant fed into the reaction side behaves like an inert gas, such as the case of oxygen permeation experiments with helium as purge. The real case would lie between these two extreme cases. Oxygen permeation flux through dense ionic or mixed conducting ceramic membranes can be related to air and reaction side oxygen partial pressures as proposed in their model. For ionic or mixed-conductors with ionic transference number close to 1 and electron conduction dominated by the transport of the electron-holes (such as yttria-stabilized zirconia and doped bismuth oxide), the oxygen permeation flux can be approximated by the following equation:JO2=k(P′hd O21/n−PΔO2 hu 1/n)  (2.11)See Bouwmeester H. J. M., Burggraaf A. J., Dense ceramic membranes for oxygen separation, in: P. J. Gellings, H. J. M. Bouwmeester (Eds.), The CRC Handbook of Solid State Electrochemistry, CRC Press, Boca Raton, Fla., 1997, Chapter 14; Han J., Xomeritakis G., Lin Y. S., Oxygen permeation through thin zirconia/yttria membranes prepared by EVD, Solid State Ionics. 1997, 93, 263-272; and Zeng Y., Lin Y. S., Oxygen permeation and oxidative coupling of methane in yttria doped bismuth oxide membrane reactor, J. Catal. 2000, 193, 58-64, each incorporated herein by reference in their entirety.
The driving force is the oxygen partial pressure in the air side minus the oxygen partial pressure in the reaction side with a positive value for constant n (i.e., n>0). For convenience in notation, this group of membranes with oxygen permeation equation in the form of Eq. (2.11) is referred to in this paper as membranes with p-type flux equation.
For ionic conductors with electron conduction dominated by the electrons, such as yttria doped zirconia at low oxygen partial pressure range, or for mixed conductors with ionic transference number close to zero, such as lanthanum cobaltite, the permeation flux equation can be expressed byJO2=k(P″O21/n—P′O21/n)  (2.12)See Park J. H., Blumenthal R. N., Electronic transport in 8 mole percent Y2O3—ZrO2, J. Electrochem. Soc. 1989, 136, 2867, incorporated herein by reference in its entirety.
The driving force is the oxygen partial pressure in the reaction side minus the oxygen partial pressure in the air side with a negative value for constant n (n<1). For convenience in notation, these groups of membranes are referred to here as membranes with n-type flux equation. Table 2.11 summarizes the parameters used. The values for constants, k and n in Eq. (2.11), for a membrane with p-type flux equation are chosen for the bismuth-based oxide Bi1.5Y0.3Sm0.2O3−δ (BYS). See Kusaba H., Shibata Y., Sasaki K., Teraoka Y., Surface (2.13) effect on oxygen permeation through dense membrane of mixed-conductive LSCF perovskite-type oxide, Solid State Ionics. 2006, vol. 177, 2249-2253, incorporated herein by reference in its entirety. The k and n values for a membrane with n-type flux equation in Eq. (2.12) are calculated from the literature oxygen permeation data given for La0.6Sr0.4Co0.2Fe0.8O3−δ (LSCF). As proposed by Xu and Thomason, The concentration of oxygen vacancies at both surfaces of the membrane (C′V and C′V) is also governed by surface exchange kinetics for the following two surface reactions (forward reaction in the feed side and backward reaction in the permeate side):
                                                        1              2                        ⁢                          O              2                                +                                    V              -                        ⁢            o                          ⁢                  ↔                                    k              f                        /                          k              r                                      ⁢                                            O              x                        ⁢            o                    +                      2            ⁢                                                  ⁢                          h              -                                                          (        B        )            where OxO represents lattice oxygen in the perovskite crystal structure, kf and kr are, respectively, the forward and reverse reaction rate constants for the forward reaction (or the reverse and forward rate constants for backward reaction). It should be noted that because of the high electronic conductivity, the electron holes are essentially constant at both surfaces of the membrane, and thus the reverse reaction rate of Reaction B and the forward reaction rate of Reaction C are pseudo zero-order at steady state under isothermal conditions. According to this, they have correlated the steady-state oxygen permeation flux as a function of P′O2, P″O2 and membrane temperature and thickness:
                              Jo          2                =                                            D              V                        ⁢                                          K                r                            ⁡                              (                                                      P                                          O                      ⁢                                                                                          ⁢                                              2                        0.5                                                              ′                                    -                                      P                                          O                      ⁢                                                                                          ⁢                                              2                        0.5                                                              ″                                                  )                                                                        2              ⁢                                                          ⁢                                                                    LK                    f                                    ⁡                                      (                                                                  P                                                  O                          ⁢                                                                                                          ⁢                          2                                                ′                                            ⁢                                              P                                                  O                          ⁢                                                                                                          ⁢                          2                                                ″                                                              )                                                  0.5                                      +                                          D                V                            ⁡                              (                                                      P                                          O                      ⁢                                                                                          ⁢                                              2                        0.5                                                              ′                                    +                                      P                                          O                      ⁢                                                                                          ⁢                                              2                        0.5                                                              ″                                                  )                                                                        (        2.13        )            Where DV, Kr and Kf are functions of temperature and the specific properties of the membrane. The values of those parameters have been determined by fitting the experimental oxygen flux data in the work done by kusaba et al. as a function of temperatures as shown in FIG. 11. FIG. 11 is a graph of the fitting of the experimental data of Oxygen permeation through LSCF1991 membrane having a thickness of 0.8 mm. As shown in the figure, both the experimental data and the present Dkk model (DV, Kr and Kf model) are in a very good agreement. Also, the as the temperature of the membrane surface increases as the oxygen permeation flux across the membrane increases due to the reduced surface resistance of the membrane as a result of temperature increase. The fitted values of the coefficients DV, Kr and Kf are listed in table 2.2 and also the values of the activation energies.
A semi-empirical form found in the literature also was used extensively to determine the local oxygen flux as a function of the membrane Temperature, the feed and permeate oxygen partial pressures, and empirical constants that depend on the specific material. This form allows for interchangeable oxygen flux mechanisms to be implemented quickly and effectively within the model in order to explore the impact of different ITM membrane materials.JO2=A*exp(−B/TM)*(P′O2n−P″O2n)  (2.14)
The pre-exponential A in some sense accounts mostly for the diffusion coefficient and the membrane thickness dependence, while B represents the effective activation energy or Arrhenius dependence due to both surface exchange kinetics and diffusion coefficient activation energy. This mechanism is chosen because it is simple, relatively accurate with respect to experimental data, and reasonably captures the impact of both surface exchange kinetics as well as the temperature dependence of the oxygen vacancy diffusion coefficient. However, it is limited in the sense that it applies for a specific membrane thickness, and also cannot be extrapolated too far from the experimental conditions used to obtain the fitted values for A and B. The functional dependence on partial pressure is assumed to be n=0.5 for LSCF, and n=0.25 for LNO based on global surface exchange kinetics theory and experimental results and the values of A and B are 26.75 mol m−2s−1pa−n, 16510 K for LSCF and 2.011 mol m−2s−1pa−n, 10240 K for LNO respectively. This is consistent with the mixed control, i.e., both diffusion and surface kinetics, in contrast to diffusion dominant where n is typically less than zero. See Mancini N. D., Mitsos A., Ion transport membrane reactors for oxy-combustione part II: analysis and comparison of alternatives, Energy. 2011, 36(8):4721-4739, incorporated herein by reference in its entirety.
Membrane Reactor Performance:
Little work was reported with a focus on examining the effects of reaction side conditions on oxygen permeation through the mixed-conducting ceramic membranes. In membrane reactors for partial oxidation of a hydrocarbon, one side of the ionic or mixed-conducting ceramic membrane is exposed to air and the other side to a hydrocarbon gas. Oxygen from the air side permeates through the membrane and reacts with the hydrocarbon in the reaction side. Although it is generally agreed that the driving force for oxygen permeation is the oxygen potential gradient across the membrane, the oxygen partial pressure, or oxygen chemical potential in the reaction side is very difficult to define or measure by experiments. See Akin F., Lin Y., Zeng Y., a comparative study on oxygen permeation and oxidative coupling of methane in disk-shaped and tubular dense membrane reactors, Ind. Eng. Chem. Res. 2001, 40, 5908-5916, incorporated herein by reference in its entirety. As a result, the quantitative value of oxygen potential gradient across an ionic or mixed-conducting ceramic membrane exposed to air and a hydrocarbon is not fixed. The permeation flux through an ionic conducting ceramic membrane depends on many parameters in a membrane reactor.
TABLE 2.1Summary of parameters used in the parametric studyParameterRangeQ0feed (ml/min)100-675T (° C.)850A (cm2)1.8k (ml/min cm2 atm1/n)BYS (ml/min cm2 atm1/n)0.035LSCF(ml/min cm2 atm1/n)0.16nBYS3.34LSCF−8.06
TABLE 2.2The obtained pre-exponential coefficients and activationenergies of DV, kf and kr for LSCF-1991 membrane fromour work through the fitting of experimental dataPre-Exponential coefficientsActivation EnergyExpressionUnitValue(kJ/mol)Dv = Dov exp(−ED/RT)m2/s1.58 × 10−573.6Kf = kof exp(−Ef/RT)m/atm0.5 · s1.11 × 1010226.9Kr = kor exp(−Er/RT)mol/m2 · s3.85 × 1011241.3
Akin and Jerry presented a simple mathematical analysis, coupled with experimental data, on the effects of hydrocarbon flow rate and reactivity with oxygen on the oxygen permeation in an ionic or mixed conducting ceramic membrane reactor for partial oxidation of hydrocarbon. In their work, Oxygen permeation experiments were conducted on a fluorite structured Bi1.5Y0.3Sm0.2O3−δ (BYS) membrane. The BYS is primarily an oxygen ionic conductor with a considerable electronic conductivity due to the doping of samarium. The membrane used in their work was a short BYS tubes, of 3-4 cm in length and 4 mm in inner diameter, were prepared from the citrate derived BYS powder followed by a green machining method. Oxygen permeation experiments were conducted, respectively, with methane and ethane as the feed to the reaction side. Oxygen permeating through the BYS membrane reacted with methane or ethane, with main reaction being oxidative coupling of methane (OCM) in the former and selective oxidation of ethane (SOE) to ethylene for the latter. The BYS membrane surface is catalytically active for OCM and SOE. Furthermore, these reactions typically involve formation of a radical on membrane surface followed by a homogeneous reaction in the gas phase inside the membrane tube. Therefore, no other catalyst was packed to the tube side of the BYS membrane in these studies. Oxygen permeation through oxygen ionic or mixed-conducting ceramic membranes under reaction conditions was analyzed in their work by a simple mathematical model considering different transport mechanisms (with n-type versus p-type flux equations) and oxidation reaction kinetics (extremely fast reaction versus no reaction). Experimental oxygen permeation data for Bi1.5Y0.3Sm0.2O (BYS) membrane under two different reaction conditions (with methane and ethane) were measured and analyzed by the modeling results.
They showed that for a membrane under reaction conditions with a specific reducing gas, the oxygen permeation flux depends strongly on the oxidation reaction rate and the reducing gas flow rate. Such dependencies are different for membranes with oxygen permeation controlled by p-type electron transport and those by n-type electron or oxygen vacancy transport. Also they reported that oxygen permeation flux under the condition of extremely fast reaction is several fold higher than that under the condition of slow reaction as a result of the decrease of the oxygen partial pressure in the reaction side with increasing reaction rate. Also they found that the oxygen permeation flux through BYS membrane under the reaction conditions with ethane is about one order of magnitude higher than that with methane because of the significantly faster oxidation reaction with ethane than with methane. The analysis results also suggest that increasing oxidation reaction rate in the reaction side by use of a heterogeneous catalyst can enhance the oxygen permeation flux through an ionic or mixed-conducting ceramic membrane. For a membrane with a fixed oxygen permeation mechanism, increasing gas flow rate lowers the conversion of oxidation reaction in the reaction side (equivalent to a decrease in reaction rate), causing a decrease in the oxygen permeation flux. Varying reactant flow rate may also cause a change in oxygen permeation mechanism leading to a complex permeation flux versus flow rate relationship.
These perovskite-type ionic conductors have technological applications in devices such as oxygen ion conducting solid electrodes, solid oxide fuel cells (SOFC), and electro-chemical oxygen pumps (EOP) and oxygen sensors. See Eng D., Stoukides M., Catalytic and electro catalyticmethane oxidation with solid oxide membranes, Catalysis Reviews Science Engineering. 1991, 33 (3 and 4), 375} 412, incorporated herein by reference in its entirety. There is also a great potential to use these materials as a membrane in a catalytic reactor. In this case, oxygen is separated from air and then used for selective oxidation reactions such as the oxidative coupling of methane (OCM) to (C2+) higher hydrocarbons or methane partial oxidation to synthesis gas. See Van Hassel B. A., ten Elshof J. E., Bouwmeester H. J. M., Oxygen permeation through La1−ySryFeO3−δ limited by carbon monoxide oxidation rate, Applied Catalaysis A. 1994, 119, 279-291, incorporated herein by reference in its entirety. Using La0.6Sr0.4Co0.2Fe0.8O3−δ (LSCF-6428) as the catalytic membrane for the OCM reaction, Xu and Thomson found that the oxygen permeation rate was higher than that required for methane activation on the membrane surface. See Xu S. J., Thomson W. J., Ion-conducting perovskite membrane for the oxidative coupling of methane, A. I. Ch. E. Journal, 1997, 44, 2731-2740, incorporated herein by reference in its entirety. Consequently, excess oxygen ions recombined to form oxygen molecules, which caused serious combustion of methane in the gas phase and reduced the selectivity of the desired C2+ products. Moreover, in a separate analysis of the stability of these membranes, they showed that morphological changes on the methane side of the membrane surface caused significant changes in the oxygen permeation flux. The surface reaction rate was found to limit the CO oxidation rate in a La1−ySryFeO3−δ (Y=0.1,0.2) membrane where the rate was nearly independent of the membrane thickness. Hence, in order to achieve practical and effective applications of these membrane reactors for partial oxidative reactions, it is essential to understand the oxygen permeation mechanism and to determine the rate limiting step among the processes of mass transfer, ionic transport, and surface reactions.
Xu and Thomson developed an explicit oxygen permeation model in a stagnation flow for ion-conducting membranes with a high ratio of electronic to ionic conductivity, which makes it possible to correlate the permeation flux to directly measurable variables. Surface exchange kinetics at each side of the membrane was emphasized and their resistance to oxygen permeation has been quantitatively distinguished from the bulk diffusion resistance. They conducted a series of experimental measurements of oxygen fluxes for La0.6Sr0.4Co0.2Fe0.8O3-δ over a wide range of temperature and oxygen partial pressures and they used the results for model regression purposes and for mechanism analysis. They concluded that the oxygen permeation at low temperatures (750° C.) is limited by the rate of oxygen-ion recombination but is dominantly controlled by bulk diffusion at high temperatures (950° C.) and this is consistent with activation energies for oxygen vacancy diffusion and for the surface exchange rates, which are estimated at 74 kJ/mol and 227, 241 kJ/mol, respectively.
The characteristics of oxyfuel combustion in an oxygen transport reactor (OTR) have been investigated. See Ben-Mansour R., Habib M., Badr H., Uddin A., Nemitallah M. A., Characteristics of Oxy-fuel Combustion in an Oxygen Transport Reactor, Energy Fuels. 2012, dx.doi.org/10.1021/ef300539c|Energy Fuels 2012, 26, 4599-4606, incorporated herein by reference in its entirety. The cylindrical reactor walls were made of dense, nonporous, mixed-conducting ceramic membranes that only allow oxygen permeation from the outside air into the combustion chamber and the simulations have been done for different composition of CH4/CO2 mixtures and for different mass flow rates. The comparison between reactive and separation-only OTR units showed that combining reaction and separation increases significantly O2 permeation rate to about 2.5 times under the assumptions given therein and the results indicated that the heat of reaction is mostly transferred to the air side with a portion used to heat the O2 permeating flux. For higher mass flow rates, the OTR operates with rich mixture resulting in low CH4 conversion.
The present disclosure describes the performance of ion transport membranes under the oxycombustion conditions in the permeate side of the membrane. Effects of many parameters including inlet gases temperature, reactor geometry, feed and sweep flow rate, oxygen partial pressure in the feed side and the effect of the percentage of CH4 in the sweep gases mixture on the permeation and combustion processes are included.
A new 3D reactor design is introduced for the substitution of ITM reactors into a gas turbine combustor including oxygen separation through ion transport membranes and then the combustion in the permeate channels with fuel in a mixture of O2/CO2/H2O. The reactor design substitutes conventional gas turbine combustor with a monolith structure ITM reactor. Particular feed and sweep flow rates are disclosed in order to meet the power required for the reactor and keeping the reactor size as compact as possible. Effects of flow configurations, channel width and percentage of CH4 in the permeate side flow are controlled under constant inlet gas temperature and fixed operating pressure of 10 bars. The reactor geometry is structured based on channel width to keep the reactor size as close as possible to the size of a conventional industrial gas turbine. The monolith structure rector design provides power ranging from 5 to 8 MWe based on cycle first law efficiency.
Based on the above details, the objectives of the present disclosure include determining experimentally and numerically, the atmospheric diffusion oxycombustion in a gas turbine model combustor, determining flame stability under wide operating range of parameters including equivalence ratio, percentage of O2/CO2 in the oxidizer mixture in addition to fuel volume flow rate, determining the minimum percentages of O2 in the oxidizer mixture required to get a stable oxy-flame, characterizing the flames to obtain the main species concentration at the exhaust section of the combustor, exhaust temperature and inside reactor temperature distribution which can serve as a database for combustion models validation, validating the numerical model using the experimental data, disclosing the oxygen permeation process and oxy-fuel combustion characteristics inside an ITM reactor using a simple symmetric design reactor allowing the reduction of the number of coordinates to 2D without reducing the accuracy of the solution, a detailed sensitivity analysis for all of the parameters affecting the membrane reactor performance, develop new oxygen permeation equation model and perform the reaction kinetics using a two-step oxy-combustion reaction kinetics model, and a new 3-D reactor design for the substitution of gas turbine combustor by an ITM reactor utilizing an ITM monolith structure reactor.